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Saturation Curve Analysis of
Humidification-Dehumidification Desalination
Systems and Analysis of Reversible Ejector
Performance OFTCHNOLOJV
by Kli M R
Ronan Killian McGovern IR IEL L-, M h'I,BE (Mechanical Engineering), University College Dublin (2010)
Submitted to the Department of Mechanical Engineeringin partial fulfillment of the requirements for the degree of
Master of Science in Mechanical Engineering
at the
MASSACHUSETTS INSTITUTE OF TECHNOLOGY
February 2012
@ Massachusetts Institute of Technology 2012. All rights reserved.
ARCHIVES
Author ..........DepartmentIof Mechanical Engineering
N' Jan 20, 2012
Certified by....John H. Lienhard V
Colli Professor of Mechanical Engineering
~ ppsis aervisor
Accepted by ................ .........................David E. Hardt
Chairman, Committee on Graduate Students
.1%
2
Saturation Curve Analysis of
Humidification-Dehumidification Desalination Systems and
Analysis of Reversible Ejector Performance
by
Ronan Killian McGovern
Submitted to the Department of Mechanical Engineeringon Jan 20, 2012, in partial fulfillment of the
requirements for the degree ofMaster of Science in Mechanical Engineering
Abstract
The limitations upon the efficiency of humidification-dehumidification (HDH) desali-
nation systems are investigated. Secondly, ejector technologies are analyzed as a
means of powering desalination systems thermally. Thermal desalination systems
offer advantages over membrane technologies in terms of robustness. They are less
prone to scaling and fouling and capable of treating feed water of higher tempera-
tures. One major drawback of thermal desalination is the high energy consumption,particularly for small scale systems. This work is motivated by the need for drasti-
cally improving the efficiency of thermal desalination technologies, particularily those
on scales of 1 to 1000 m3 per day.Given multiple liquid and gas streams and simultaneous heat and mass transfer,
the limits on the performance of HDH systems are not easily recognisable. By con-
sidering the saturation curve for moist air, a methodology is proposed to graphically
understand these limits. A pinch point methodology is employed to evaluate the im-
pact of heat and mass exchanger size upon performance. The impact of salinity upon
the design methodology is described. As the pinch point temperature and concen-
tration differences decrease, the gained output ratio (GOR) is shown to increase and
recovery ratio is shown also to increase. There is shown to be a critical pinch point
temperature difference beyond which there is no advantage to having multiple stages,each with different liquid to dry air mass flow ratios. Finally, the GOR is shown to
increase as the temperature range of the cycle decreases.Little attention is given in the literature to the definition of an ejector efficiency.
Here, the entrainment ratio of real devices is compared to the reversible entrain-
ment ratio and denoted the reversible entrainment ratio efficiency. The definition
of different performance metrics for ejectors are analyzed and compared. Graphical
illustrations are provided to support intuitive understanding of these metrics. The
performance metrics are applied to existing experimental data. The behaviour of
reversible ejectors is investigated. Analytical equations are also formulated for re-
versible ideal-gas ejectors. For general air-air and steam-steam ejectors, the exergeticefficiency rx is found to be very close in numerical value to the reversible entrainmentratio efficiency, r/RER.
Thesis Supervisor: John H. Lienhard VTitle: Collins Professor of Mechanical Engineering
Acknowledgments
I would like to express my great appreciation for the support provided by the Interna-
tional Fulbright Science & Technology Award funded through the U.S. Department
of State. I would like to thank the King Fahd University of Petroleum and Minerals
for funding this research through the Center for Clean Water and Clean Energy at
MIT and KFUPM.
Ba mhaith liom buiochas a gabhail do mo theaghlach, Jim, Stephanie, Aoife,
Fergal agus an cailin a bhi chomh-lach liom 6n li gur shroich m6 an tir seo, Ivonne.
Tdim anseo de bharr bhur tacaiochta. Beidh me ar ais libh go mion minic. Go raibh
mile maith agaibh.
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Contents
1 Introduction 15
1.1 Understanding the Limits upon the Energy Efficiency of HDH . . . . 16
1.2 Characterizing Ejector Performance . . . . . . . . . . . . . . . . . . . 18
2 Evaluation of the Performance Limits of Humidification Dehumidi-
fication Desalination Systems via a Saturation Curve Analysis 21
Nom enclature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23
2.1 M otivation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
2.2 Performance Metrics . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
2.3 Mass Balances, Enthalpy Balances and Saturation Curves . . . . . . . 30
2.3.1 Mass Balance and the Saturation Curve . . . . . . . . . . . . 30
2.3.2 Energy Balance and the Saturation Curve . . . . . . . . . . . 32
2.4 Application of the Saturation Curve Method . . . . . . . . . . . . . . 35
2.4.1 Modelling Approximations . . . . . . . . . . . . . . . . . . . . 35
2.4.2 Zero Extraction Cycle . . . . . . . . . . . . . . . . . . . . . . 38
2.4.3 Single Extraction/Injection Cycle . . . . . . . . . . . . . . . . 40
2.5 Results: Explanation of the Performance Limits of HDH . . . . . . . 42
2.5.1 Impact of a Single Extraction/Injection upon GOR and RR 44
2.5.2 Effect of the Cycle Temperature Range . . . . . . . . . . . . . 47
2.5.3 Effect of Component Size . . . . . . . . . . . . . . . . . . . . . 49
2.5.4 Impact of Salinity upon HDH Design and Performance . . . . 54
2.6 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57
3 Analysis of Reversible Ejectors and Definition of an Ejector Effi-
ciency 59
Nom enclature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 60
3.1 Introduction.. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 62
3.2 Description of Reversible Ejector Processes . . . . . . . . . . . . . . . 63
3.2.1 Equations Describing a Thermodynamically Reversible Process 64
3.2.2 Definition of Thermodynamically Reversible Reference Processes 66
3.2.3 Interpretation of a Thermodynamically Reversible Ejector Pro-
cess . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 70
3.3 Reversible Entrainment Ratio Calculations for Different Fluids . . . . 71
3.3.1 A ir-air . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 72
3.3.2 Steam-steam Reversible Entrainment Ratios . . . . . . . . . . 74
3.3.3 Steam-Moist-Gas Reversible Entrainment Ratios . . . . . . . . 81
3.4 Development of Ejector Performance Metrics . . . . . . . . . . . . . . 83
3.4.1 Reversible Entrainment Ratio Efficiency . . . . . . . . . . . . 87
3.4.2 Reversible Discharge Pressure Efficiency . . . . . . . . . . . . 88
3.4.3 Turbine-Compressor Efficiency . . . . . . . . . . . . . . . . . . 88
3.4.4 Compression Efficiency . . . . . . . . . . . . . . . . . . . . . . 90
3.4.5 Exergetic Efficiency . . . . . . . . . . . . . . . . . . . . . . . . 91
3.5 Comparison of Ejector Performance Metrics using Experimental Data 93
3.6 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97
4 Recommendations 99
4.1 Reduced Temperature Range HDH Cycles . . . . . . . . . . . . . . . 99
4.2 Future Research Directions . . . . . . . . . . . . . . . . . . . . . . . . 100
4.3 Ejector-Expander Power Generation System . . . . . . . . . . . . . . 101
List of Figures
1-1 Thermal energy requirements versus scale of thermally driven desali-
nation technologies . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16
1-2 Humidification-dehumidification process . . . . . . . . . . . . . . . . 17
1-3 Ejector Driven HDH System . . . . . . . . . . . . . . . . . . . . . . . 19
2-1 Water heated, open air, humidification dehumidification desalination
system . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . 26
2-2 Control volume for a general humidification or dehumidification process 31
2-3 Molar ratio of vapor to carrier gas versus saturation temperature . . . 32
2-4 Saturation curve on an enthalpy-temperature diagram for a moist car-
rier gas on an enthalpy versus temperature diagram . . . . . . . . . . 34
2-5 Saturation pressure of moist air when at equilibrium with pure water
and seawater . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37
2-6 Zero extraction process paths - zero pinch point temperature differences 39
2-7 Single extraction process paths - zero pinch point temperature differences 41
2-8 Process paths of liquid and moist air streams within a zero extrac-
tion/injection HDH system . . . . . . . . . . . . . . . . . . . . . . . . 42
2-9 Saturation humidity ratio at the moist air bulk temperature and the
feed temperature for zero and single extractions . . . . . . . . . . . . 46
2-10 Effect of the top moist air temperature upon GOR and RR, Tf,DB =
Tma,H,B = 250C . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47
2-11 Illustration of the effect of decreasing top air temperature upon the
recovery ratio . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48
2-12 Effect of the inlet feed and saturated moist air inlet temperature upon
GOR and RR, Tma,D,T = Tma,H,T = . . . . . . - - -.. . . . 49
2-13 Zero extraction process paths - PPTD = 2 K . . . . . . . . . . . . . . 51
2-14 Single extraction process paths - PPTD = 2 K . . . . . . . . . . . . . 52
2-15 GOR versus temperature pinch within the humidifier and dehumidifier
for cases with zero and a single water extraction . . . . . . . . . . . . 53
2-16 Evaporating water temperature in the humidifier at the high tempera-
ture pinch point, the water extraction point and the low temperature
pinch point - all versus PPTD in the humidifier and dehumidifier . . 54
2-17 Recovery ratio versus temperature pinch within the humidifier and
dehumidifier for cases with zero and a single water extraction/injection 55
3-1 Schematic diagram of a steady flow ejector . . . . . . . . . . . . . . . 62
3-2 Control volume for an ejector process . . . . . . . . . . . . . . . . . . 64
3-3 Graphical representation of the reversible entrainment ratio . . . . . . 66
3-4 Graphical representation of the reversible discharge pressure reference
process . . . . . . . .. . . ............................ 68
3-5 Graphical representation of the reversible entrainment ratio reference
process. . . . . . . . .. . . ... . .. . . . ................ 69
3-6 Schematic diagram of a thermodynamically reversible ejector for fluids
of identical chemical composition . . . . . . . . . . . . . . . . . . . . 71
3-7 Process paths for a thermodynamically reversible ejector . . . . . . . 72
3-8 Lines of constant ratio of specific heats, -y, for steam . . . . . . . . . . 75
3-9 Specific enthalpy-entropy diagram for pure water . . . . . . . . . . . 76
3-10 Reversible entrainment ratio versus motive steam pressure for a steam-
steam ejector. (PD=15 kPa, PE=10 kPa, xM=l, XE=1) . . . . . . . . 77
3-11 Optimal motive steam pressure of a steam-steam ejector . . . . . . . 78
3-12 Reversible entrainment ratio versus discharge pressure for a steam-
steam ejector. (PM=10 bar, PE=10 kPa, xM=l, XE=1 . . . . . . .. 79
3-13 Reversible entrainment ratio versus motive steam superheat for a steam-
steam ejector. (PD=15 kPa, Pm= 10 bar, PE=10 kPa, XE=l, TIsat =
1800 C ......... ................................... 79
3-14 Reversible entrainment ratio versus motive steam specific entropy for
a steam-steam ejector, with motive steam at constant temperature.
(TNI=179.9 0 C, Psat,M=10 bar, PD=15 kPa, PE-10 kPa, XE=1) . . . - 80
3-15 Isotherms on a specific enthalpy-entropy diagram for pure water . . . 81
3-16 Reversible entrainment ratio versus motive steam pressure for steam-
moist-carrier-gas ejectors. (PD=75 kPa, PE=50 kPa, TE=50 C, xM=1,
E1) . .. .... ...... ....... ....... ....... ... 84
3-17 Reversible vapor entrainment ratio versus motive steam pressure for
steam-moist-carrier-gas ejectors. (Uc= 1.5, PE=50 kPa, TE=50 C, XM=1,
#E=1) . . . . . . . .-.. . . . . . ..-....... ................. 84
3-18 Reversible vapor entrainment ratio versus entrained saturated carrier
gas temperature for a steam-moist-carrier-gas ejector. (PD=75 kPa,
PE=50 kPa, Ui=20, XM=1, #E=1) . . - . - - - - - - . . . . . . . . . 85
3-19 Isentropic expansion and compression processes . . . . . . . . . . . . 89
3-20 Comparison of efficiency definitions for a steam-steam ejector over a
range of compression ratios . . . . . . . . . . . . . . . . . . . . . . . . 94
3-21 Comparison of efficiency definitions using data from an air-air ejector,
[1, 2]. Chamber throat diameter, 0.88 inches; Diffuser exit diameter,
3.02 inches; Nozzle throat diameter, 0.5625 inches; Nozzle exit area
0.735 inches; Plenum diameter for secondary flow, 3.25 inches. .... 95
3-22 Comparison of efficiency definitions using data from a steam-steam
ejector operating at conditions of critical back pressure [3]. Chamber
throat diameter, 18 mm; Diffuser exit diameter, 40 mm; Nozzle throat
diameter, 2 mm; Nozzle exit area 8 mm; Plenum diameter for secondary
flow, 24 mm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 96
4-1 Low temperature range HDH configuration . . . . . . . . . . . . . . . 100
4-2 Ejector-expander shaft-power generation system....... . . ... 102
List of Tables
3.1 Reversible discharge pressure process . . . . . . . . . . . . . . . . . . 67
3.2 Reversible entrainment ratio process . . . . . . . . . . . . . . . . . . 68
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Chapter 1
Introduction
Evaporation is one of the simplest ways one can conceive of a water purification
system. Compared to membrane technologies such as reverse osmosis, thermally
driven desalination technologies offer many advantages. Thermal systems are less
prone to scaling and fouling and have the ability to deal with feed water of higher
temperatures. A range of thermally driven desalination technologies is indicated in
Fig. 1-1.
At the lower end of system scale in Fig. 1-1 lie solar stills, the simplest but
most inefficient means of desalination. At the top end lie multi-effect distillation
and multi-stage flash, systems that exist only at large scales and that exhibit higher
efficiency. Between these two extremes lies a gap. To some extent, this gap is filled
by energy efficient electrically driven technologies such as reverse osmosis (RO) or
mechanical vapor compression (MVC). However, feed water of high salinity (beyond
approximately 70 000 ppm) requires hydraulic pressures beyond what RO membranes
can tolerate. The concentration of feed water achievable using mechanical vapor
compression is also limited due to scaling concerns. RO is also unsuitable for high
temperature feed water (beyond approximately 400 C) due to membrane deterioration.
Furthermore, both RO and MVC cannot be driven by heat, without the provision
of a heat engine. In remote environments with high levels of insolation, photovoltaic
panels rather than solar heaters are required.
Returning to the gap in Fig. 1-1, we can ask whether this may be the place for
104
0 Solar Still
102 _H Multi-Stage Flash
101 Multi-Effect Distillation10
100 I I I I I I I I I I r
10~1 100 101 106System Scale [m3/day]
Figure 1-1: Thermal energy requirements versus scale of thermally driven desalinationtechnologies
humidification-dehumidification (HDH) desalination, a technology almost as simple
as a solar still but with the potential for higher efficiency. As it currently stands, it
may be argued that low energy efficiency is the limiting factor in the deployment of
HDH, particularly if the system is to be driven by solar power. With thermal energy
requirements on the order of 100 kWh/m 3, the area and cost of solar collectors is
prohibitive. Following this line of argument, for the HDH technology to progress,
improvements in the energy efficiency of the system and the cost of providing that
energy are required.
1.1 Understanding the Limits upon the Energy Ef-
ficiency of HDH
The HDH system, as outlined in Fig. 1-2 is not a simple one to analyze. These
processes involve three distinct streams, feed water, moist carrier gas and condensate
intercommunicating through simultaneous heat and mass transfer. Air enters the
humidifier and is heated and humidified by a counterflow of hot water. The moist hot
Hot H Heater
Moist Air Feed
Humidifier DehumidifierProduct
Water
Inlet j Brine Feed Outletair i Reject Water + air
Figure 1-2: Humidification-dehumidification process
air is subsequently condensed in the dehumidifier as heat is transferred to the feed
water. The pre-heated feed water at the outlet of the dehumidifier further increases
in temperature in the water heater before being sprayed into the humidifier.
The optimal design of an HDH system requires the careful selection of the mass
flow rate and temperature of each stream within the system. Given that the humid-
ifier and the dehumidifier are intrinsically coupled, such optimal design requires an
intuitive and complete understanding of the system's operation. Returning to Fig. 1-
1, we can summarise the motivation for this research by posing the two following
questions:
1. Why is the energy efficiency of HDH lower than other thermally driven tech-
nologies?
2. Through an improved understanding of the system and enhanced design method-
ologies, what are ultimately the limits upon energetic efficiency?
1.2 Characterizing Ejector Performance
Beyond the improvements in HDH system performance, sought in the Chapter 2 of
this thesis, we can also consider means of reducing the cost of the thermal energy
input to desalination systems. For example, at the simplest level, we could ask
whether it is best to use solar energy for heating water or generating steam. Higher
temperature solar collectors provide a higher exergetic input although their capital
cost may be higher. For the same quantity of heat collected, the available work is
greater. However, in the context of thermal desalination we must be vigilant about
the risks of scaling at higher temperatures. We would like a heat source of high
temperature but would like to maintain low brine temperatures. Using ejectors as
a means of introducing a thermal energy input to a system is one solution to this
apparent paradox.
An ejector, as indicated within Fig. 1-3, accelerates high pressure vapor to high
velocity and low pressure. This low pressure provides a suction force that draws in a
secondary fluid stream. After mixing, the fluid is decelerated in a diffuser and reaches
the discharge pressure. Rather than solely providing a heating effect, an ejector allows
flow work to be done on the entrained stream. Although the steam enters at high
temperature, streams within the desalination system can be maintained at lower
temperatures, thus avoiding the risks of scaling.
The role of the ejector in Fig. 1-3 is to provide a difference in pressure between the
humidifier and dehumidifier, an idea that was initiated and developed by Narayan et
al. [4, 5].
Beyond the application of Fig. 1-3, ejectors are widely employed in multi-effect-
distillation desalination plants. Their use in vapor compression systems as a replace-
ment for the mechanical compressor has also been suggested. In simple terms, an
ejector serves as a mechanical compressor but at a much lower cost, due to its simple
design with no moving parts. However, it is well known that the efficiency of ejectors
is low, which limits their viability, particularly where solar collectors are concerned.
Interestingly, when we posed the question as to how efficiency could be improved, we
i";Ap*uuri ' Power Out
Figure 1-3: Ejector Driven HDH System
found that the definition itself of efficiency was not rigorously defined or investigated
in literature. Before adopting an approach to improve ejector performance it became
necessary to define appropriate performance parameters. These observations led to
the two key questions motivating this ejector research.
1. How can a metric of ejector performance be defined by benchmarking real ejector
processes against reversible processes?
2. How can the behaviour of a thermodynamically reversible ejector be described
mathematically and graphically?
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Chapter 2
Evaluation of the Performance
Limits of Humidification
Dehumidification Desalination
Systems via a Saturation Curve
Analysis
Ronan K. McGovern, Gregory P. Thiel, Syed M. Zubair, John H.
Lienhard V
Abstract
Due to the challenges involved in modelling combined heat and mass transfer pro-cesses, analyses of humidification dehumidification systems are typically experimentalor numerical. Few provide an intuitive insight into the performance of HDH systemsin terms of energetic efficiency (Gained Output Ratio - GOR) and recovery ratio.This chapter aims to bridge that gap. A methodology, revolving around the sat-uration curve of a moist carrier gas, is developed to analyze HDH systems. Theproblem of choosing optimal liquid and carrier gas mass flow rates within the systemis reduced to a geometric problem involving graphs of enthalpy versus temperature.By employing a single extraction/injection of liquid from the humidifier to the dehu-midfier the energetic performance is improved due to better matching between thesaturation curve and the feed streams on an enthalpy-temperature diagram. The
GOR is shown to increase as the temperature range of the cycle decreases. A pinchpoint methodology is employed to evaluate the impact of heat and mass exchangersize upon performance. As the pinch point temperature and concentration differencesdecrease, the GOR is shown to increase and recovery ratio is shown also to increase.There is shown to be a critical pinch point temperature difference beyond which thereis no advantage to having multiple stages, each with different liquid to dry air massflow ratios. Finally, the impact of salinity upon the design methodology is described.
Nomenclature
Acronyms
HDH Humdification-Dehumidification
GOR Gained Output Ratio
PPTD Pinch Point Temperature Difference
PCTD Pinch Point Concentration Difference
RR Recovery Ratio
Symbols
c concentration [mol/m 31
c, specific heat capacity at constant pressure [kJ/kg-K]
h specific enthalpy [kJ/kg]
hfg latent heat of vaporization [kJ/kg]
H enthalpy [kJ]
mass flux [kg/m 2 -s]
rh mass flow rate [kg/s]
M molar mass [kg/kmol]
N moles
p partial pressure [bar or kPa]
P absolute pressure [bar or kPa]
Q heat input [kJ]
T temperature [C or K]
w specific work [kJ/kg product]
Z compressibility factor
Greek
A change
# relative humidity [-]
x molar humidity [mol H20/mol carrier gas]
w humidity ratio [kg H20/kg carrier gas]
Subscripts
B bottom
cg carrier gas
da dry air
D dehumidifier
E external
f feed
H humidifier
in into system
1 liquid
ma moist air
mm minimum
p product
r reject
T top
v vapor
Superscripts
molar
rate [s-1]
* per unit dry carrier gas
2.1 Motivation
Humidification-dehumidification (HDH) systems are commonly viewed as robust liq-
uid purification systems, driven by heat and capable of operating on a small scale de-
centralized basis. Energy consumption is high compared to membrane based reverse
osmosis technologies [6] but the components are simple and require low maintenance.
In desalination applications, HDH systems purify saline feed water. From another
perspective, the role of HDH systems could be to concentrate industrial waste water
streams.
The premise underlying HDH is that by humidifying air (or another suitable gas)
through the evaporation of a pure substance from a feed solution and its subsequent
condensation, pure water vapor may be separated from its solutes and suspended
solids. Subsequently, in the dehumidification process, the vapor is condensed to
form a purified liquid product. The HDH system, Fig. 2-1, which separates the
humidification and dehumidification processes, is one means of implementing a carrier
gas based liquid purification process. This is in contrast to dewvaporation systems,
where heat is transfered from the dehumidifier to the humidifier via a common heat
transfer wall, [7].
The system described in Fig. 2-1 is a water heated open air HDH process. Air
enters the humidifier from the surroundings and is heated and humidified by a coun-
terflow of feed. The moist air then flows to the dehumidifier, in which the air is cooled
and water vapor is condensed by a counterflow of liquid feed. The feed is heated in
the dehumidifier and in the heater before being cooled and partially evaporated in
the humidifier. The feed remaining at the bottom of the humidifier, known as the
reject, exits the system. Crucial to the energy efficiency of carrier gas based liquid
purification processes is the recovery of heat during air cooling and dehumidification
for use during air humidification and heating.
Although HDH systems are potentially simple in terms of design and operation,
their poor energy efficiency is a significant drawback. Some of the best systems
still require 150 kWh of thermal energy per m3 product water [8]. This inefficiency
MMAAB Mi=,D.B
Figure 2-1: Water heated, open air, humidification dehumidification desalination sys-tem
has led to to numerous analytical, numerical and especially experimental studies [9].
The prevalance of experimental studies is somewhat of a testament to the difficulties
involved in analyzing a system with three fluid streams (feed, condensate and carrier
gas) undergoing simultaneous heat and mass transfer. At least two approaches to
the analytical modelling of HDH have been taken. The first, employed by Narayan
et al. [101, involves the use of an enthalpy based effectiveness to quantify humidifier
and dehumidifier performance in a parametric study of HDH systems. Narayan et
al. consider the thermal energy requirements for fixed component effectivenesses
and cycle temperature range. By achieving a heat capacity rate ratio (liquid to
moist air stream) of unity in the dehumidifier, the entropy generation per unit water
produced in the system is minimized. In papers by Hou et al. [11], Hou [12] and
Zamen et al. [13], the performance of the humidifier and dehumidifier is parametrized
by specifying pinch point temperature differences. A key idea in this work is that
the process path of moist air is considered to follow the saturation curve. Hou et
al. [11] demonstrate that the thermal energy recovery rates increase rapidly as the
temperature differences at the pinch points within the system decrease. Hou [12]
later explains that, by having a two stage cycle and therefore reducing the relative
mass flow of air to water in the top of the cycle, the curvature of the saturation curve
can be reduced. Thus, temperature differences between streams at the inlet and exit
of the dehumidifer can be reduced. Zamen et al. [13] further build upon the work
of Chou and describe a multi-stage HDH cycle within which the mass flow of air is
reduced within each stage of higher temperature. This further flattens the saturation
curve. The thermal energy requirements of the cycle are shown to decrease with an
increasing number of stages.
Given this overview of analytical studies, we can list some of the key remaining
questions:
1. How does the choice of the cycle temperature range affect thermal energy re-
quirements?
2. How does the improvement achieved by implementing a multi-stage system
relate to component sizes (e.g. pinch point temperature difference)?
3. What are the limits upon the fraction of feed water recoverable as product?
4. How does feed salinity affect the thermal energy requirements of HDH desali-
nation?
The goal of this chapter is to address each of these questions in an intuitive manner,
building upon and unifying results found in literature. Specifically, this chapter will
describe a method of analysis, involving the saturation curve of a moist carrier gas,
to provide a graphical insight into the design of HDH systems. In Section 2.2, the
performance metrics for HDH systems are outlined. In Section 2.3, the saturation
curves for a moist carrier gas and their relevance to the analysis of HDH are outlined.
The saturation curve methodology is applied in Section 2.4. Finally, in Sections 2.5
and 2.6, the answers to the motivating questions above are provided.
2.2 Performance Metrics
Before proceeding with our discussion of approximations, we define key performance
parameters for the HDH process. The first parameter of interest concerns the ener-
getic performance of the system, measured by considering the mass (or moles) of water
produced per unit of heat input to the cycle. Since this quantity is dimensional, it is
common to multiply by the latent heat of vaporisation to obtain the Gained Output
Ratio, or the GOR, a dimensionless quantity.
GOR= .hfg _Nphf (2.1)Q Q
The GOR is a measure of the latent heat of water produced per unit of heat input.
There is no consensus in the literature upon the temperature at which the latent heat
of vaporization should be computed. However, the current authors would suggest that
it be evaluated at ambient temperature. In this way, we can envision a system that
evaporates water at ambient temperature by operating at reduced pressure and sub-
sequently condensing the water vapor without recovery of the heat of condensation.
Such a system would have a GOR of 1.
Many HDH systems operate with a constant mass (or molar) flow rate of carrier
gas. It is therefore useful to define all quantities on a carrier gas basis, i.e. per
unit mass of carrier gas or per mole of carrier gas. Employing an asterisk to iden-
tify quantities measured on a carrier gas basis, the following expression for GOR is
obtained:
Ashfg AxhfgGOR= .I - . (2.2)
Q* Q*Then for a water heated system we have:
Ashrg AXhfgGOR= = ~ (2.3)
Ah* Ah*
The second measure of interest concerns the fraction of feed water that the system
is capable of recovering as product. The recovery ratio captures this fraction and is
defined as the ratio of product water to feed water, either on a mass or molar basis:
RR= (2.4)mf,H20 Nf,H20
The final measure of performance is the exergetic efficiency. This performance
metric allows us to determine how close system performance is to being thermody-
namically reversible. The exergetic efficiency is defined as the useful exergetic output
of the system divided by the exergetic input. A detailed discussion of this efficiency
for desalination processes is provided by Mistry et al. [6]:
711 = "= ( " -n) (2.5)Xin n) i
wmin refers to to the minimum work required to extract one unit of pure water from an
infinite body of seawater. For seawater with a salinity of 35,000 ppm, the minimum
work (calculated with feed and product streams at 25*)C is 2.71 kJ/kg pure water
produced. For a water heated HDH cycle, the exergetic input to the system is in
the form of heat transfer to the feed. Tij is the temperature at which heat enters
the control volume of the HDH system. This temperature is somewhat ambiguous.
Strictly speaking, if the feed is heated from 65'C to 70*C, for example, heat would
enter the system over a range of temperatures. However, since the temperature range
within the heater is small, heat will be considered to enter at a constant temperature,
equal to the top temperature of the feed in the cycle. For an HDH cycle, it is
convenient to write Eq. (2.5) in terms of GOR:
711 =GOR wrlIf (2.6)
hy, I1 - T
Of course, if analyzing an HDH system with pure water as the feed stream, the
system provides no useful exergetic output, since pure water is produced from a feed
of pure water. The exergetic efficiency is not of relevance in analyzing such a process.
In other words, the second law efficiency only becomes useful when there is a useful
exergetic output, i.e. when a saline feed stream is being purified.
2.3 Mass Balances, Enthalpy Balances and Satu-
ration Curves
Saturation curves are extremely useful in visualising mass flows and energy flows
within HDH. A plot of humidity ratio versus temperature allows us to visualise mass
transfer, whilst a plot of enthalpy versus temperature illustrates energy transfer.
These two plots share a common trait, in that their y-axes (humidity ratio and specific
enthalpy) are both defined on a unit mass basis. This is possible since the mass
flow rate of air through the system is constant. In this section, we describe how
the conservation of mass and energy within humidifiers and dehumidifiers can be
graphically represented. We also describe the influence of salinity upon the saturation
curves.
2.3.1 Mass Balance and the Saturation Curve
In Fig. 2-2, we consider a general control volume for humidification or dehumidifica-
tion. A stream of liquid and a stream of moist air enter and exit the control volume.
Individual control volumes are also indicated for the gas mixture and the liquid. A
portion of each individual control volume passes just inside the liquid on the liquid
side of the gas-vapor interface.
When the mass flow of air is constant throughout the system, we may specify all other
mass flow rates and state variables on a unit dry air basis. Conducting a vapor mass
balance upon the overall control volume and the individual control volumes yields the
following relations:
drih = -Thadw = jvdA (2.7)
For dehumidification, Eq. (2.7) describes the relation between the condensate pro-
duction and the change in humidity ratio of the air. For humidification, Eq. (2.7)
relates the mass of feed evaporated to the change in humidity ratio of the air. To quan-
tify the rate of condensation or the rate of evaporation as a function of air temperature
rh co rn o+do)
h,, haCVa
cv
k, ha ha+dh, h,,+dh,
interface, i
dAQ .... .
/I, h+dhj
Figure 2-2: Control volume for a general humidification or dehumidification process
we can consider a graph of the saturated humidity ratio versus temperature. Two
saturation curves for an ideal carrier gas are provided in Fig. 2-3. They correspond to
a saturated ideal carrier gas in equilibrium with pure water and in equilibrium with
seawater (at a salinity of 35 000 ppm). The properties moist air mixture in equilib-
rium with pure water are evaluated using the formulations presented by Hyland and
Wexler [14]. The moist air mixture in equilibrium with seawater is approximated as
an ideal mixture in calculating its properties and the saturation pressure of seawater
is computed using correlations developed by Sharqawy et al. [15].
If the carrier gas is ideal, then the saturation curve on a molar basis is independent of
the choice of carrier gas. Commonly in psychrometric charts the humidity is displayed
on a mass basis, in which case Fig. 2-3 would be scaled by the molar mass ratio of
the carrier gas. The ability of the ideal carrier gas to hold vapor increases rapidly
with temperature due to the exponential dependence of vapor pressure on saturation
temperature.
During dehumidification of a carrier gas-vapor mixture, the relevant saturation
curve will always be that which is in equilibrium with pure water. During humidifica-
tion, salinity will influence the vapor pressure and thus the saturation humidity ratio
of the carrier gas. The difference in value of the saturation humidity ratio between
-Pure Water
- - Seawater
0.4
00.3
0.2
-0.1
0.060 70 80
Figure 2-3: Molar ratio of vapor to carrier gas versus saturation temperature
seawater and pure water systems is approximately -8% at 25*C and -4% at 700C.
2.3.2 Energy Balance and the Saturation Curve
Returning to Fig. 2-2, and applying the First Law to the control volume of carrier
gas, we may relate the heat transfer rate to the specific enthalpies and mass flow
rates:
- cQ = Tha(dha + wdhv) + rhahvdW + jvhi,idA (2.8)
Employing Eq. (2.7), we substitute for j,.
- 6O = lla [dha + wdh, + (hv - hi,j)dw] (2.9)
6 QE refers to an external heat transfer to the liquid stream. In an externally adiabatic
humidifier this heat transfer would be zero. In a dehumidifier QE would be negative
and would represent heat transfer to a coolant. Applying the First Law to the liquid
) 20 30 40 50Temperature [*C]
control volume yields:
SQE -+- = ~ rdhl + hldrh - jvhidA (2.10)h*/h*
ri d71lW(E + SQ dh1 + . (h, - h,) (2.11)
ma ma
The above equations give rise to the definition of specific enthalpy on a dry air basis,
denoted by h*, to be used in all further analyses. For dehumidification, h* would
represent the specific enthalpy of the condensate on a dry air basis. In the case of
humidification, h* would represent the specific enthalpy of the feed on a dry air basis.
Eq. (2.10) indicates that the process paths of liquid streams can be easily visualised
on a diagram of h* versus T. In the humidifier the evaporation rate is small compared
to the mass flow of liquid. Liquid lines appear straight if drawn on Fig. 2-4 since c,,1
is almost constant with temperature. The slope of such a line will be dictated by the
mass flow ratio (or the molar flow ratio) of liquid to dry air. In Fig. 2-4 the enthalpy
per mole of dry air, h*, is plotted versus the saturation temperature, for the cases
where moist air is in equilibrium with pure water and with seawater (at 35 000 ppm).
Again, during dehumidification of a carrier gas-vapor mixture, the relevant saturation
curve will always be that which is in equilibrium with water. During humidification,
salinity will influence the vapor pressure and thus the saturation specific enthalpy of
the moist air. The difference in value of the specific enthalpy of moist air between
seawater and pure water systems is approximately -2% at 25*C and -3% at 70*C.
We would like to obtain an expression describing the shape of the saturation curve
(equilibrium with pure water) in Fig. 2-4. Returning to Eq. (2.9), we can consider the
order of magnitudes of each of the terms in dh*,. We begin by considering h, - hi,j.
hv - hi,; ~ c,, [T, - Tt(P,)] + hfg(P,) + c, [Tsat(Pv) - T,j] (2.12)
~hg(Pv) (2.13)
25,000
20,000 -I
-Pure Watere 15,000 -
- - Seawater
10,000 -9
5,000
0 -
0
-5,00010 20 30 40 50 60 70 80
Temperature [*C]
Figure 2-4: Saturation curve on an enthalpy-temperature diagram for a moist carriergas on an enthalpy versus temperature diagram
The air and the vapor are at the same temperature, Ta, and the temperature difference
between the liquid and the air would typically be below 10 *C. Consequently, hfg
dominates Eq. (2.12). We now introduce this approximation into Eq. (2.9) to obtain
Eq. (2.14)
dwdh* ~iia (cp,a + wcp,,v + hfg ) dTa (2.14)
dh* ~ hahfgdW (2.15)
The humidity ratio of saturated air at 50*C is approximately 0.1 kg vap/kg air, while
the slope of the saturation curve dw/dTa is approximately 0.005. At moderate and
high temperatures (approximately 40 to 100*C) the value of dh* is primarily dictated
by the third term in (2.14). Hence, the shape of the saturation curve on a humidity-
temperature and an enthalpy-temperature diagram are very similar in this region.
This assertion will prove useful when evaluating the performance of HDH cycles in
Sect. 2.5.
2.4 Application of the Saturation Curve Method
2.4.1 Modelling Approximations
The following four approximations are employed within our analysis:
1. Constant Pressure
2. Ideal Gas Behavior
3. Zero Salinity Feed
4. Saturation Curve Process Path
We now deal with each of these in turn.
Constant Pressure Approximation
Embodiments of the HDH cycle do exist within which the pressure of the gas mix-
ture is varied in order to drive humidification and dehumidification, [4, 5]. In this
work, only systems in which the pressure is approximately constant throughout are
considered. Of course, a pressure gradient must be provided by a fan to circulate the
carrier gas. Typically the pressure rise provided by fan would be expected to be on
the order of a few kPa, compared to an absolute pressure of 100 kPa. Whilst this
requires work input to the system, this work is small compared to the heat input to
the system.
Ideal Gas Approximation
For HVAC analyses, the ideal gas approximation is very well accepted when modelling
moist air mixtures. Sincee HDH systems operate over a higher range of temperature,
and may be designed employing gases other than air, it is worth verifying the ac-
curacy of ideal gas approximations. HDH systems operate at temperatures between
approximately 20*C and 80*C and at pressures of 1 bar or lower, [161. The ideality
of gases may be verified by considering the compressibility factor at typical pressures
and temperatures of HDH systems. A compressibility factor close to unity indicates
an ideal gas:
z PVPt- (2.16)RT
Calculation of this compressibility factor for helium, air, carbon dioxide and vapor
indicates that, for the range of temperatures and partial pressures experienced in
HDH systems, the compressibility factor is always between 0.99 and 1. For example,
the compressibility factor of water vapor, helium, air and carbon dioxide in saturated
vapor-carrier gas mixtures in thermal equilibrium with pure water at 50*C are 0.9961,
1, 0.9999 and 0.9967, to the nearest four decimal places. Whilst the properties of
steam are known to deviate from ideal when close to saturation, the temperatures
and pressures present in HDH systems are sufficiently low for this deviation to be
negligible.
Effects of Salinity
It would be incorrect to state that the effect of salinity upon the design of HDH
systems is negligible. First of all, saline feed solutions carry a risk of scaling that can
greatly affect system operation and performance. However, even beyond the effects
of scaling, salinity affects the energetic performance of HDH in two key ways:
1. Salinity increases the specific heat at constant pressure of water.
2. Salinity lowers the the saturation pressure of vapor.
Using the seawater correlations developed by Sharqawy et al. [15], we find that for a
range of temperatures between 20*C and 80*C, the specific heat capacity of seawater
at 35 ppm on a mass basis, is approximately 4 to 4.5% smaller than that of pure
water. Salinity also causes a decrease in the saturation pressure of water. In Fig. 2-5,
the saturation pressure is plotted for pure water and water within a seawater solution
of 35 000 ppm. The difference in saturation pressure (seawater to pure water) is
approximately -2% over the temperature range of interest for HDH. This will become
important when evaluating the impact of salinity upon the shape of the saturation
curve on humidity ratio and enthalpy diagrams in Section 2.3.
35
30
~25
20
. 15
10
5
0
-Pure Water
- - Seawater
10 20 30 40 50Temperature [*C]
Figure 2-5: Saturation pressureseawater
60 70 80
of moist air when at equilibrium with pure water and
The effects of salinity upon the specific heat of brine and the saturation humidity
of air are less than 5%. The methodology involved in designing an energy efficient
HDH system based upon a pure feed stream is the same as if the stream were to be
saline. As such, the consideration of a pure water HDH system is a useful, simpler
and yet reasonably accurate means of developing a design methodology.
Saturation Curve Process Path
The saturation curve methodology for the analysis of HDH systems involves approx-
imating the process path of the moist carrier gas during humidification or dehumid-
ification as the saturation curve. At saturation, the temperature and vapor pressure
are not independent. Consequently, the parametric space that the moist carrier gas
may occupy is reduced to the saturation curve.
Making the above four approximations simplifies the analysis of HDH considerably.
Leaving salinity aside, the state of the water at any point in the system is fully
specified by its temperature. The state of the gas mixture is also fully specified by its
temperature, since the pressure of the system is constant, and the saturation curve
relates the humidity ratio to temperature. Next we describe two types of water heated
open air cycles, the first without any extraction or injection and the second with one
single extraction and injection of water (as in Fig. 2-1).
2.4.2 Zero Extraction Cycle
We begin by describing the the process paths of the moist air, condensate and feed
streams within the dehumidifier. We fix the temperature of the inlet feed stream at
Tf,D,B, and also fix the top temperature of the moist air, Tma,T. Next we imagine
a dehumidifier of infinite area such that the pinch point temperature difference at
the top and bottom are zero. In other words Tf,D,B = Tma,D,B = Tc,D,B and also
Tf,DT = Tma,D,T. By fixing the temperature of the moist air at any point, we are also
fixing its humidity, since we are approximating the process path by the saturation
curve. By making these specifications, the mass flow ratio of feed to dry air becomes
fixed, in order for the First Law to be satisfied for the dehumidifier. This may be
seen in Fig. 2-6. At each point in the dehumidifier, the condensate is taken to be
at the temperature of the saturated moist air. The dashed line in Fig. 2-6b includes
the enthalpy of the moist air and of the condensate. The choice of whether the
condensate is at the liquid or air temperature is insignificant since the effect of the
change in enthalpy of the condensate upon the shape of the dashed line is very small.
As outlined in Eq. (2.14), the shape of the saturation curve is dictated by the vapor
content of the air. Since the dehumidifier is externally adiabatic, the change in
enthalpy of the feed water matches the change in enthalpy of the moist air and the
condensate exactly.
Since the cycle is water heated, the top temperature of the moist air is the same as
the top temperature of the dehumidifier. Let us fix the inlet air temperature, Tma,H,B-
We are now left with two unknowns, Tma,H,T and Tf,H,B, and one equation, the First
Law for the externally adiabatic humidifier:
800
700
600
500
400
300
200
100
0
-100
800
700
600
500
400
300
200
100
0
-100
feed water
80 100
'2moist air andcondensate
0 20 40Temperature [*C]
(b) Dehumidifier
Figure 2-6: Zero extraction process paths - zero pinch point temperature differences
0 20 40 60Temperature ["C]
(a) Humidifier
h~mar - haHB = hf,H,T - hr,H,B
However, by setting a pinch point temperature difference of zero in the humidifier,
we obtain a final relation that closes the system of equations. At the pinch point in
the humidifier, the evaporating water line must be tangent to the saturation curve.
The process path of the evaporating water is close to but not exactly a straight line.
The path is slightly convex due to the evaporation of water. Since the mass of water
evaporated is small the curvature is not noticeable.
2.4.3 Single Extraction/Injection Cycle
Again, for this single extraction/injection cycle, we begin by considering the dehu-
midifier. We fix the inlet feed stream at Tf,D,B and the top temperature of the moist
air, Tma,T. We set the pinch point temperature difference to be zero at the top and
bottom of the dehumidifier but also at the point of injection, i.e. Tf,M,D = Tma,M,D,
Fig. 2-7b.
At this stage we have two equations, the First Law above and below the injection
point. We also have three unknowns, the temperature of injection, Tf,M,D, and the
feed to dry air mass flow ratio above and below the point of injection. Moving to
the humidifier, we again fix the inlet air temperature, Tma,H,B. We note that the feed
temperature at the extraction point in the humidifier equals the feed temperature
at the injection point in the dehumidifier. The mass flow ratio at the top of the
humidifier is set by the mass flow ratio above the injection point in the dehumidifier.
The mass flow ratio just below the point of extraction in the humidifier is set by the
mass flow ratio below the injection point in the dehumidifier, Fig. 2-7b. We are now
left with four unknown temperatures, Tma,H,T, Tf,H,B, Tf,M,D and the temperature of
the moist air at the feed extraction point in the humidifier, Tma,H,M. The mass flow
ratio above and below the injection point in the dehumidifier are also unknown. This
leaves us with six unknowns. However, we can also find six equations. First, we have
four equations from the application of the first law above and below the extraction
moist air
extraction
evaporatingwater
800 -
700 -
600 -
500 -
9400 -
*n 300 -
S200 -
100
0
-100-0
feed water
0 20 40Temperature [*C]
(b) Dehumidifier
60 80
\ waterinjection
moist air andcondensate
60 80
Figure 2-7: Single extraction process paths - zero pinch point temperature differences
20 40Temperature [*C]
(a) Humidifier
800
700
600
500
400
300
200
100
0
-100
- -
- -
point in the humidifier and above and below the injection point in the dehumidifier.
Additionally, by setting pinch point temperature differences of zero in the humidifier,
we have two further equations. As in the case with zero extractions/injections, the
slope of the evaporating stream must equal the slope of the saturation curve in the
humidifier at the pinch points. This allows us to solve for the profiles obtained in
Fig. 2-7.
2.5 Results: Explanation of the Performance Lim-
its of HDH
In this section we return to the questions posed in Section. 2.1 and provide graphical
answers. We begin by explaining how our performance metrics, GOR and RR, relate
to the graphs presented in Section 2.4. To help with these explanations, it is instruc-
tive to combine Fig. 2-6 to form Figs. 2-8a and 2-8b for a zero extraction/injection
system.
heateroutlet
0 00
feed water
T- -o f 6Vevaporating
stream
20 30 40 50 60 70 80Temperature [*C]
Figure 2-8: Process paths oftion/injection HDH system
liquid and moist air streams within a zero extrac-
Although Fig. 2-8 is approximative it will prove to be very illustrative in explaining
trends in the performance of HDH cycles. Feed water enters the dehumidifier and is
pre-heated. The moist air, in counterflow, is cooled and water vapor is condensed to
form a pure product. The feed water increases in temperature within the heater. The
feed is then cooled and water evaporates into the counterflow of moist air, which is
heated and humidified. In this case, air enters the system at 30 0C and feed enters at
27*C. The top moist air temperature is approximately 60 0C.
In HDH, the latent heat required for vaporization in the humidifier is solely pro-
vided by the sensible cooling of the feed stream. A large mass flow rate of feed is
required per unit of water evaporated. Thus, the water produced in an HDH system
per unit of feed, in other words the recovery ratio, is low. Strictly speaking, due to
evaporation, the ratio of liquid to moist air flow rate in the humidifier is not con-
stant. However, since the mass of water evaporated is low, the liquid path (that is
slightly curved in Fig. 2-6a) is well approximated by a straight line in Fig. 2-8. In
Fig. 2-6b the change in enthalpy of the condensate stream has been incorporated into
the curve representing changes in enthalpy of the moist air (since both streams are
approximated as being at the same temperature). However, since the mass flow of
condensate per unit air is small, there is little difference between the dashed lines in
Fig. 2-6a and Fig. 2-6b. In Fig. 2-8 the saturation curve shown represents the path
of the moist air in the humidifier and the dehumidifier.
The saturation curve on an enthalpy-temperature diagram is approximately the
same as the saturation curve on an humidity-temperature diagram, scaled by hfg.
Therefore, the height of the saturation curve in Fig. 2-8 is proportional to the quantity
of water produced per unit of dry air. Meanwhile, the ratio of feed to dry air flow
is given by the slope of the liquid lines in Fig. 2-8, which may be represented by
the height of the curve divided by the width. Consequently, the recovery ratio is
proportional to the width of the saturation curve in Fig. 2-8.
Returning to our definition of GOR, Eq. (2.3), we are reminded that the heat
input per unit mass of air is given by the change in enthalpy of the feed stream in the
heater per unit mass of air, Eq. (2.18). Combining Eqs. (2.3) and (2.18) we obtain
FffH ,T
Eq. (2.19). This change in enthalpy is easily visualized in Fig. 2-8, where '". Cp,f is
the slope of the feed line. The feed process path within the heater may be visualised
by extending the feed line from the heater inlet to the heater outlet temperature. The
further apart the heater inlet and heater outlet temperature are, the greater the heat
input to the cycle.
h*,D,T - hf,H,T fHT cp,f(Tf,D,T - Tf,H,T) (2-18mra
GOR = M Awhfg (2.19)mhT Cp,f (Tf,D,T - Tf,H,T)
2.5.1 Impact of a Single Extraction/Injection upon GOR and
RR
Consider the water produced per unit of carrier gas. The top temperature of the moist
air is fixed at 70*. Due to our choice of pinch point temperature differences of zero
in the humidifier and dehumidifier, the temperature reached by the saturated moist
air at the exit of the dehumidifier equals the inlet feed temperature. Consequently,
the water produced per unit mass (or mole) of carrier gas in the dehumidifier is the
same for the zero and single extraction systems.
The implementation of a single extraction results in an improvement of the ener-
getic performance (GOR) simply because it allows a smaller temperature difference
across the heater. Also, if the temperature difference across the heater is reduced,
the temperature difference between the rejected stream and the feed stream will also
reduce according to a First Law energy balance upon the system. For the single
extraction system of Fig. 2-7, the GOR is 14 compared to the zero extraction system
of Fig. 2-6, where the GOR is 3.5.
In essence, by employing an extraction/injection, the process paths of the feed
provide a better piecewise linear approximation of the saturation curve. From a
second law perspective we are really saying that by improving the matching of the
saturation curve and the liquid lines we are distributing the driving forces for heat
and mass transfer throughout the system in a more even fashion, see [17] and [18].
This reduces the stream to stream temperature differences in the humidifier and
dehumidifier, thus minimizing the heat requirements of the cycle. Comparing Fig. 2-
7 to Fig. 2-6, we can see that the employment of a single extraction greatly reduces the
average driving force for heat transfer (the stream to stream temperature difference).
Mass transfer is also important process within HDH, driven by differences in
vapor concentration. At the liquid-air interface in the humidifier, and the heat-
exchanger-surface-air interface in the dehumidifier, the air is locally saturated at
a temperature close to that of the the liquid, assuming that air is the dominant
heat transfer resistance. By plotting the bulk humidity of the air stream versus the
saturation humidity at the liquid temperature, we may visualize the concentration
difference that drives mass transfer in the humidifier and the dehumidifier, Fig. 2-9.
From Fig. 2-9 it is clear that the implementation of an extraction/injection greatly
reduces the average driving force for mass transfer (concentration difference) within
both the humidifier and the dehumidifier. In Fig. 2-9a it is interesting to note the
very large difference in humidity ratio at the top of the humidifier. This is because a
small temperature difference between the bulk air and the liquid at high temperature
results in a large difference in humidity ratio due to the steep slope of the saturation
curve at high temperatures.
Finally, we turn to the recovery ratio, RR. We have already said that for fixed
top and bottom moist air temperatures, the water produced per unit of carrier gas is
fixed. Now, the recovery ratio is given by the mass of product divided by the mass
of feed, or in other terms, the mass of product per unit mass of dry air divided by
the mass of feed per unit mass of dry air, Eq. (2.20). The mass flow rate of feed per
unit mass of carrier gas at the inlet to the HDH systems in Fig. 2-6b and Fig. 2-7b
is given by the slope of the feed line at the inlet. Clearly the slope is lower in the
system with one single extraction and consequently the recovery ratio is higher.
RR = rP -. damf,H20 mnda Mf,H20
10 30 50Temperature [*C]
(a) Humidifier
moist air andcondensate
' (si leextraction)
10 30 50Temperature [*C
(b) Dehumidifier
Figure 2-9: Saturation humidity ratio at the moist air bulktemperature for zero and single extractions
temperature and the feed
-C 0.6
0.5
0.4
0.3
0.2
0.1
0
0.3
08 0.2
0.15
0.1
0.05
0
70 90
feed(zero
extraction)
70 90
2.5.2 Effect of the Cycle Temperature Range
By temperature range, we mean the difference in the top moist air temperature in the
cycle and the bottom moist air temperature (i.e. at the humidifier inlet). In general,
the moist air inlet temperature will be fixed by the temperature of the surroundings
whilst the top moist air temperature is a design parameter. In Fig. 2-10 the effect of
the top moist air temperature upon the GOR and the RR is plotted.
45 - - 12
single 'extraction/
mction 0135 -
30 - 8
/ zero extraction/
1/--
30 40 50 60 70 80Top moist air temperature [*C]
Figure 2-10: Effect of the top moist air temperature upon GOR and RR, Tf,D,B
Tma,H,B = 25*C
As the top moist air temperature decreases, the GOR increases. Simply said, this
is because our piecewise linear approximation of the saturation curve becomes more
accurate when considering a cycle that spans a smaller temperature range. Fig. 2-10
also indicates that the RR decreases as the top moist air temperature is lowered. This
is because the mass flow of feed per unit mass of air (dictated by the slope of the
feed line at the inlet, see Fig. 2-6b) decreases at a faster rate than the mass flow of
product water per unit mass of air. To make this intuitively obvious, we may consider
Fig. 2-11.
According to Eq. (2.14), the saturation curves on an humidity-temperature and on
evaporatingstream
- b
feed stream/ I ||a a
20 30 40 50 60 70 80
Temperature [*C]
Figure 2-11: Illustration of the effect of decreasing top air temperature upon the
recovery ratio
an enthalpy-temperature diagram are scaled copies of each other. Therefore, distance
b on Fig. 2-11 is proportional to the change in humidity of the carrier gas in the
dehumidifier,7 and consequently the water produced per unit of air. As the top moist
a2
air temperature decreases, the length of b decreases and the length of a decreases 2
Since the slope of the feed line is given by k, the rate of water production per unit
mass of air, given by b, decreases at a faster rate than the ratio of feed to air flow,
given by k. Hence, the recovery ratio must decrease, according to Eq. (2.20).
aa
In Fig. 2-12 we investigate the effect of increasing the feed and moist air inlet
temperatures upon the GOR and RR, for a fixed top moist air temperature. As the
inlet streams' increase in temperature, the water produced decreases slowly. (Due
to the exponential shape of saturation curve the quantity of water produced is much
more sensitive to the top than the bottom temperature.) Meanwhile, the heat input
per unit of air decreases more rapidly, resulting in an improvement of the energetic
performance and the GOR. Applying similar arguments to those outlined through
2Note that the feed line and the evaporating line must be parallel as the liquid to air mass flowratios are equal in the humidifier and dehumidifier
25 -10
20 8
P4 15 - + -- 6
10 zero extraction/ _ 4 einjection
5 _ 2
0 020 25 30 35 40 45 50 55
Inlet feed and moist air temperature [*C]
Figure 2-12: Effect of the inlet feed and saturated moist air inlet temperature uponGOR and RR, Tma,D,T = Tma,H,T = 70C
the use of Fig. 2-11, it may be shown that the RR must also decrease as the bottom
inlet streams' tempeature increases.
2.5.3 Effect of Component Size
In a bi-fluid heat exchanger, the driving force for heat transfer is dictated by tem-
perature differences between the two streams. For a fixed duty, i.e. heat transfer
rate, a smaller heat exchanger will require larger driving force (assuming constant
heat transfer coefficients). The pinch point temperature difference (PPTD) is defined
as the minimum temperature difference between the two streams. It is somewhat
characteristic of the driving force for heat transfer. Following this line of argument,
to achieve a fixed rate of heat transfer within a heat exchanger, the pinch point tem-
perature difference is somewhat indicative of the heat exchanger size required. As
the PPTD increases, a smaller heat exchanger is required and vice versa.
The above arguments may also be applied for a mass exchanger. The driving force
for mass transfer is given by a concentration difference. A pinch point concentration
difference (PPCD) would be somewhat analogous to the PPTD in a heat exchanger.
In a heat and mass exchanger (HMX), such as an humidifier or dehumidifier, there
are two driving forces, temperature difference and concentration difference. The size
of the heat and mass exchanger, for a fixed heat and mass transfer rate is therefore
somewhat related to the PPTD and the PPCD. In the case of the saturation curve
method of HDH analysis, the temperature and humidity of the moist air are always
linked by the equation describing the saturation curve. Hence, by specifying a PPTD
between a fluid and moist air stream, we are in fact also specifying a PPCD between
the fluid and the moist air stream.
We know that the saturation vapor concentration is an exponential function of
temperature. Therefore, the PPCD and PPTD are not linearly related. In fact,
depending upon the temperature at which we specify the PPTD, the PPCD could be
very small (at low temperature) and very large (at high temperature). In other words,
by increasing the PPTD it is unclear to what extent the heat and mass exchanger
size is decreasing. What we can say, is that for fixed top and bottom moist air
temperatures, increasing the PPTD in the humidifier or dehumidifier will decrease
the HMX size required for a fixed water output.
For illustrative purposes a graphical representation of HDH cycles with zero and
a single extraction/injection for pinch point temperature differences of 2 K in the
humidifier and dehumidifier are provided in Fig. 2-13 and also Fig. 2-14. Note how,
due to the shape of the saturation curve, there are pinch points at the inlet and outlet
of the dehumidifier but within the humidifier.
In Fig. 2-15 the GOR is plotted versus PPTD, setting the temperature difference at
each pinch point to be equal at all pinch points. Again the inlet streams' temperature
is 25*C and the top moist air temperature is 70*C.
Unsurprisingly, as the PPTD increases, the GOR decreases due to an increasing
temperature difference across the heater. Interestingly, there exists a critical value of
PPTD beyond which there is no advantage in employing an extraction and injection
of water from the humidifier to the dehumidifier. This effect may be explained by
considering a plot of the feed temperatures at which the pinch points occur in the
moist air
pinch =2 K
800
700
600
500
9400
300
200
100
0
-1000 20 40 60
Temperature [*C]
(a) Humidifier
80 100
pinch = 2 K
feed water
/moist air andcondensate
pinch = 2 K
0 20 40Temperature [*C]
(b) Dehumidifier
60 80
Figure 2-13: Zero extraction process paths - PPTD = 2 K
evaporatingwater
800
700
600
500PC
400
300
200
100
0
-100
800
700
600
500
-400
300
200
100
0
-100
800
700
600
500
'400
' 300
S200
100
0
-100
pinch = 2 K
moist air /
pinch = 2K
0 20 40Temperature [*C]
(a) Humidifier
pinch = 2K
water .injection
feed water /
pinch = 2 K
Figure 2-14: Single extraction process paths - PPTD = 2 K
waterextraction
water
60 80
condensate
0 20 40Temperature [*C]
(b) Dehumidifier
60 80
14
12
10 -
98 - singleextraction/
extractionsl2 -ijinoctions2
00 2 4 6 8 10
Pinch point temperature difference [K]
Figure 2-15: GOR versus temperature pinch within the humidifier and dehumidifierfor cases with zero and a single water extraction
humidifier and the extraction temperature versus the PPTD, Fig. 2-16.
As the PPTD increases, the optimal extraction temperature increases. Once the
top pinch point reaches the top moist air temperature, no further advantage may be
obtained via extraction/injection.
A second observation arising from Fig. 2-15 is that the GOR of a system with a
single extraction/injection is more sensitive to changes in PPTD than a system with
zero extractions. In a system with zero extractions and a small PPTD, Fig. 2-13, the
PPTD is not a very good measure of average driving forces within the system. Rather,
the average driving force is dictated by mismatch in shape between the saturation
curve and the feed streams. Increasing the PPTD therefore has a weak effect upon
the temperature difference obtained across the heater. By contrast, in Fig. 2-14a and
Fig. 2-14b, the PPTD is much more representative of the mean temperature difference
between the streams. This is because much of the mismatch in shape between the
liquid and moist air paths has been eliminated by the single extraction/injection.
Consequently, the temperature difference across the heater is much more sensitive to
80 -
top pinch
70 - point
-------------------------------
60 -------- extraction!
injection pointU 50 -.*6--
- -- bottom pinch
40 point
930
20 -
10 -
00 1 2 3
Pinch point temperature difference [K]
Figure 2-16: Evaporating water temperature in the humidifier at the high temperaturepinch point, the water extraction point and the low temperature pinch point - allversus PPTD in the humidifier and dehumidifier
the PPTD.
Finally, in Fig. 2-17 we plot the RR versus the PPTD. As with the GOR, the RR
decreases with increasing PPTD. As PPTD increases, the moist air temperature at
the outlet of the dehumidifier increases, as does its humidity. Consequently, the mass
flow rate of water produced per unit mass of air decreases. Additionally, the slope
of the liquid lines must increase, meaning that the mass flow rate of feed per unit
mass of air must increase. Thus, the mass of water produced per mass of feed must
decrease, again according to Eq (2.20).
2.5.4 Impact of Salinity upon HDH Design and Performance
The above analysis of an HDH cycle, done neglecting the effects of salinity, gives an
intuitive insight into the coupling between the humidifier and dehumidifier. At this
stage, it is worth pointing out how the analysis would have differed had seawater
properties been taken into account.
singleextraction/
10 -\ injection
8 - %zero
o %extractions/injections.
6
4
2-
00 2 4 6 8 10
Pinch point temperature difference [K]
Figure 2-17: Recovery ratio versus temperature pinch within the humidifier and de-humidifier for cases with zero and a single water extraction/injection
The first difference is the smaller value of specific heat at constant pressure of the
feed water in a system with saline feed. At a salinity of 35 000 ppm, the specific
heat of seawater is about 4.5% smaller than that of pure water. If the same feed to
air mass flow ratio was chosen for a system with such saline feed, the slope of the
feed stream in the dehumidifier and the evaporating stream in the humidifier would
decrease by about 4.5%. To counteract this, and to achieve the same matching of the
process paths in the humidifier and dehumidifier, the mass flow ratio of feed to dry
air would have to be increased by about 4.5%. By adjusting the mass flow rate the
product of mass flow rate and specific heat capacity is held constant and this effect
of salinity is largely negated.
Seawater salinity will reduce the slope of the saturation curve within the humidi-
fier, due to the lower vapor pressure of seawater compared to pure water at the same
saturation temperature. If the same PPTD is to be maintained as in the system with
pure feed, the evaporating feed line in Fig. 2-8 will shift to the right, resulting in a
larger temperature difference across the heater.
In a system with saline feed water, the properties of the evaporating stream within
the humidifier will vary due to changing salinity. However, since the recovery ratio
of HDH cycles is very low (less than 12% of the feed water is evaporated and sub-
sequently condensed), the change of salinity due to concentration of the brine in the
humidifier will be very small. In other words, if we consider the effect of salinity
upon specific heat to be a second order effect (4.5%), then the effect of changing
salinity due to evaporation would be a third order effect. In a system with a single
extraction/injection, the salinity of the brine introduced from the humidifier to the
dehumidifier at the midpoint would be slightly different from the salinity of the feed
just before the dehumidifier injection point.
2.6 Conclusions
The coupling between the humidifier and dehumidifier makes analysis of HDH system
performance challenging. In this chapter, we have developed a graphical method that
enables evaluation of the thermal performance (GOR) and water recovery of an open
air HDH system. By limiting the parametric space that may be occupied by the state
of a moist carrier gas to the saturation curve, the analysis is simplified. This method
is applicable to cycles both with and without extraction/injection from the humidifier
to the dehumidifier. The main conclusions arising from this chapter are as follows:
1. A graphical method of analysis has been developed, which facilitates the exact
assessment of the role the top and bottom moist air temperatures and the feed
to air mass flow rate ratio upon system performance.
2. This method also facilitates the exact assessment of the role of injection and
extraction in improving the thermal performance of an HDH system.
3. As the temperature range of an HDH cycle increases, the GOR decreases since
the piecewise linear approximation, by the liquid streams, of the exponential
saturation curve is a poorer match.
4. As the pinch point temperature difference increases, the GOR and the recovery
ratio decrease.
5. The recovery ratio of an HDH system is limited by the fact that a large mass
flow of hot brine is required at the humidifier inlet to provide the sensible heat
necessary to supply the latent heat of vaporization.
6. A single extraction/injection improves energetic performance (GOR) by improv-
ing the matching of shape between the saturation curve and the quasi-linear feed
streams.
7. There exists a critical pinch point temperature difference beyond which there
is no advantage to extracting and injecting water from the humidifier into the
dehumidifier.
THIS PAGE INTENTIONALLY LEFT BLANK
Chapter 3
Analysis of Reversible Ejectors and
Definition of an Ejector Efficiency
Ronan K. McGovern, G. Prakash Narayan, John H. Lienhard V
Abstract
Second Law analyses of ejector performance have rarely been conducted in litera-ture. Many measures of ejector efficiency have not always been clearly defined andthe rationale underlying and justifying current performance metrics is often unclear.One common means of assessing performance is to define a thermodynamically re-versible reference process against which real processes may be benchmarked. Thesereversible processes represent the thermodynamic limit of real ejector performance.In this chapter, parameters from real and reversible processes are compared and per-formance metrics are defined. In particular, the entrainment ratio of real devicesis compared to the reversible entrainment ratio and denoted the reversible entrain-ment ratio efficiency. An efficiency comparing the ejector performance to that of aturbine-compressor system is also investigated, as is an exergetic efficiency. A rigorous
analysis of performance metrics reported in the literature is undertaken. Graphicalillustrations are provided to support intuitive understanding of these metrics. Ana-
lytical equations are also formulated for ideal-gas models. The performance metrics
are then applied to existing experimental data to illustrate the difference in theirnumerical values. The reversible entrainment ratio efficiency rE R is shown to al-
ways be lower than the turbine-compressor efficiency r/rER. For general air-air and
steam-steam ejectors, the exergetic efficiency qx is very close in numerical value to
the reversible entrainment ratio efficiency, IRER-
Nomenclature
Symbols
c, specific heat capacity at constant pressure [kJ/kg.K]
ER entrainment ratio [-]
h specific enthalpy [kJ/kg]
h specific molar enthalpy [kJ/kmol]
hE,D specific enthalpy of the entrained fluid at the discharge [kJ/kg]
hM,D specific enthalpy of the motive fluid at the discharge [kJ/kg]
j dissipation [J]
i mass flow rate [kg/s]
M molar mass [kg/kmol]
$ molar flow rate [kmol/s]
p partial pressure [bar or kPa]
P absolute pressure [bar or kPa]
Q rate of heat transfer [W]
R gas constant [kJ/kg.K]
RER Reversible Entrainment Ratio [-]
s specific entropy [kJ/kg.K]
9 specific molar entropy [kJ/kmol-K]
Sgen entropy generation rate [kW/K]
9gen dimensionless entropy generation [-]
T temperature [C or K]
v specific volume [m3/kg]
x dryness fraction of steam [-] or specific exergy [J/kg]
y useful work done [J]
Greek
7 ratio of specific heats
A change
r; efficiency [-]
Ulc compression ratio (PD/PE) [-I
# relative humidity [-]
W absolute humidity [kg H20/kg carrier gas]
ratio of moles of vapor to moles carrier gas [-]
Subscripts
C compression
cg carrier gas
comp compressor
D discharge fluid
E entrained fluid
f saturated liquid
g saturated vapour
irrev associated with irreversible dissipation
M motive fluid
NE nozzle exit
q associated with heat transfer
RDP reversible discharge pressure
RER reversible entrainment ratio
RHE reversible heat engine
s isentropic
TER turbine-compressor entrainment ratio
turb turbine
vap vapor
X exergetic
0 ambient
Superscripts
rev reversible
sat saturated
Discharge
Motive NE Fluid
Fluid M D
tEEntrained Fluid
Figure 3-1: Schematic diagram of a steady flow ejector
3.1 Introduction
Steady flow ejectors are devices without moving parts that typically combine two
fluids, one of high total pressure and one of low total pressure, to obtain a fluid at
an intermediate total pressure at the discharge. The fluid with high total pressure,
known as the motive fluid, is expanded through a nozzle to high velocity and low
pressure (point NE in Fig.3-1). The low pressure at the nozzle exit provides the
driving force for suction of the low pressure fluid, known as the entrained fluid, at
pressure PE, such that Pm > PE > PNE. The motive and entrained streams undergo
mixing, brought about by shear forces between the fluids, before the resulting mixture
is diffused to attain the discharge pressure, PD, Fig. 3-1.
The most intuitive although approximative manner to comprehend ejector opera-
tion is to analyse performance using simple 1-D analyses. Examples of excellent work
in this regard have been done by Elrod [19], Keenan et al. [20], Chou et al. [21] and
Arbel et al. [22]. Ejectors are relatively simple in design and can provide reliable
operation at low capital cost. These advantages have led to ejectors being used in
processes requiring the drawing of a vacuum (e.g. extraction of non-condensible gases
from condensers) or the compression of a gas (e.g. within refrigeration cycles). In
thermal desalination, ejectors are commonly employed to compress and entrain water
vapor. Such devices are known as thermal vapor compressors. The increased use
of thermal vapor compressors for energy intensive processes of thermal desalination,
in particular multi-effect distillation [23-26], has brought about renewed interest in
the efficiency of ejectors. Despite the interest in energy efficiency, scant attention
has been paid to the definition and justification of the utility of ejector performance
metrics employed to date. Bulusu et al. [1] presented common measures of ejector per-
formance, whilst Arbel et al. [22] performed an analysis of entropy generation within
ejectors. Nonetheless, there remains no universally accepted efficiency to characterise
and compare the performance of such devices.
In this chapter, ejector performance is evaluated by benchmarking real ejector
processes against carefully defined thermodynamically reversible processes. In Sec-
tion 3.2, a systematic methodology is presented to define such processes. One impor-
tant parameter arising from this analysis is the reversible entrainment ratio, i.e. the
maximum theoretical entrainment ratio achievable by a thermodynamically reversible
ejector. The variation in the reversible entrainment ratio with an ejector's operating
conditions is investigated in Section 3.3. Then, in Section 3.4 ejector performance
metrics are explained and justified by means of analytical equations or graphical
methods. Finally, in Section 3.5, the numerical values of these metrics are computed
for experimental data available in open literature. The exergetic efficiency -rx of an
ideal gas ejector with inlet fluids at the same temperature is shown to be identical to
the reversible entrainment ratio efficiency r/R-
3.2 Description of Reversible Ejector Processes
This section establishes a basis for benchmarking real device performance against the
performance of ideal, thermodynamically reversible processes. In order to effectuate
this comparison, the following steps are necessary:
1. Identify the input quantities, output quantities and equations describing a ther-
modynamically reversible process.
2. Define a thermodynamically reversible reference process, against which real pro-
cesses may be compared.
Motive Fluidrhm,PM,hmsm Discharge
Fluid
M rnD, PhD SDControl Volume
L --------- ----- JEntrained Fluid
rhE,P,hE SE
Figure 3-2: Control volume for an ejector process
3. Develop a performance metric based on a comparison of parameters between
the real and reversible processes. (3.4)
A control volume model of an ejector process is presented in Fig. 3-2. The inlet
stream of higher pressure is known as the motive flow, the low pressure inlet stream
is termed the entrained flow and the outlet stream is termed the discharge flow. The
control volume in both real and reversible cases is considered externally adiabatic
(Qo = 0). Due to the high flow speeds within ejectors, the rate of heat loss from the
device to ambient is taken to be small relative to changes of enthalpy of the streams
within. The nomenclature of Fig. 3-2 shall henceforth be used in refering to the mass
flow rates and state properties of inlet and outlet streams. All analyses shall consider
only equilibrium states of inlet and outlet fluids where the fluids are at conditions
of stagnation (zero kinetic energy). Thus, all temperatures and pressures are total
rather than static temperature and pressures.
3.2.1 Equations Describing a Thermodynamically Reversible
Process
Taking a control volume approach, the First and Second Laws of Thermodynamics
may be written for a real ejector process involving pure and identical fluids at the
ejector inlet:
First Law of Thermodynamics for an adiabatic process:
rnMhM +rnEhE = ThDhD (3.1)
hM + ER - hE = (1 + ER)hD (3.2)
The entrainment ratio, ER is defined as the ratio of entrained mass flow to motive
mass flow,
ER = ME (3.3)7m
Second Law of Thermodynamics for an adiabatic process:
rhMSM + TnESE + Sgen = rnDSD (3.4)
sM + ER -SE -Sgen = (1 + ER)SD (3.5)mM
Noting that no entropy is generated during a reversible process, the equations for an
adiabatic and reversible process may be formulated as follows:
hM + RER - hE = (1 + RER)hrv (3.6)
SM + RER - SE = (1 + RER)s'v (3.7)
Reversible entrainment ratio RER is defined as the ratio of entrained mass flow to
motive mass flow in the reversible process:
RER= (E rev (3.8)
For a pure fluid, the thermodynamic state is fully specified at the control volume
inlets and outlet once two state properties (i.e. h and s) are known. An equation
h RER RER+1
RER +1
PD =fn(sD,hD)
RER
PT
0 RER+I
Specific Entropy [J/kg-KI S
Figure 3-3: Graphical representation of the reversible entrainment ratio
of state may be written relating the temperature T or the pressure P to the specific
enthalpy and entropy. The specific enthalpy and entropy of the discharge stream
is simply the mass weighted average of the inlet streams, see Fig. 3-3. Of course,
according to Eq. (3.6) and (3.7), the entrainment ratio, ER, is the mass ratio relevant
to this weighting. Hence, on a diagram of specific enthalpy versus specific entropy,
the discharge fluid state must lie on the line joining the states of the inlet fluids, as
shown in Fig. 3-3. The next step is to define the different possible reversible reference
processes.
3.2.2 Definition of Thermodynamically Reversible Reference
Processes
Six state variables and one mass flow rate ratio describe the process of Fig. 3-2. For
a real process there are two equations, Eq. (3.2) and Eq. (3.5), including one further
unknown, S.ger. This value of this entropy generation term is dictated by the quality
of design of the device within the control volume. For a reversible process, there are
two associated equations, Eq. (3.6) and Eq. (3.7). Consequently, given 7 variables
and 2 equations, there are five degrees of freedom that must be specified in order
to define a reversible reference process. There are therefore 21 (7-choose-5) different
ways in which a reference process could be defined. A select number of combinations
are discussed in the following subsections.
Reversible Discharge Pressure Reference Process
In a paper elucidating the sources of irreversibilty within steady flow ejectors, Arbel
et al. [22] investigated ejector performance in great detail by comparing real processes
to a reversible reference process, described in Table 3.1:
Table 3.1: Reversible discharge pressure process
Fixed Output
PE = P PTE =rev rev
PM = a
TM = Tr
ER = RER
In essence, Arbel et al.'s reference processes asks: for fixed inlet fluid states and
a fixed entrainment ratio, what is the maximum discharge pressure achievable if the
process is reversible? On an h - s diagram, Arbel's process may be represented as:
Discharge enthalpy in both the real and reference case is a weighted average of
the inlet fluid enthalpies. Since by definition of this reference process, the real and
reversible entrainment ratios are equal, the discharge specific enthalpy is also equal in
the real and reversible cases, Eq. (3.2). In the real process, entropy is generated and
thus the discharge specific entropy is greater than the weighted mean of the entropy
of the inlet streams.
For applications where the inlet states are fixed, as is the relative mass flow of
motive and entrained fluid available, this reference process may be employed to de-
termine how close the actual discharge pressure achieved is to the reversible limit.
RER ERRER+1 ER+1
1 1
RER+1 ER+1
Specific Entropy [J/kg-K] S
Figure 3-4: Graphical representation of the reversible discharge pressure referenceprocess
Reversible Entrainment Ratio Reference Process
In many situations, the inlet states to the ejector are known, as is a desired discharge
pressure1 . Such is the case when one has steam available at a fixed state and wishes
to compress vapor, also at a known state, to a desired pressure. The efficiency of the
ejector will determine the mass flow of vapor one can entrain per unit mass flow of
steam (i.e. the entrainment ratio). In such a scenario, the reference process described
in Table 3.2 is of relevance:
Table 3.2: Reversible entrainment ratio process
Fixed Output
PE = ev RER
TE = Tev ev
PM = PT
TM = M'
PD = PD*V
'Where one is dealing with an ejector involving fluids of identical chemical composition, it isequivalent to define either the discharge pressure or the saturation temperature at that pressure, asis done by Al Khalidy and Zayonia [27] and by Elrod [19].
h
P;' = PD
PM, TMER
-s IER+1 RER
% %IC%
E RER+1
ER+1
Specific Entropy [J/kg-K] S
Figure 3-5: Graphical representation of the reversible entrainment ratio referenceprocess
On an h - s diagram, the reversible entrainment ratio reference process may be
represented as:
The real discharge pressure in Fig. 3-5 is the same as that in Fig. 3-4. However,
the discharge state is different. In the reversible entrainment ratio reference process,
the reversible and real discharge pressure are set as equal. Employing the reversible
entrainment ratio reference process, one may compare the real entrainment ratio to
the entrainment ratio achieved in a reversible process.
Summary and Further Reference Processes
The above two reference processes are only two of the 21 possible reference processes
for a dual inlet, pure and identical fluid ejector. In particular, knowledge of the
entrainment ratio is interesting from a cost perspective as it quantifies the mass
of fluid entrained per unit mass of motive fluid supplied. The reversible discharge
pressure reference process is also considered due to the notable work presented on
this process by Arbel et al. [22]. One could also consider a reference process whereby
the motive and discharge state are fixed along with the entrainment ratio, and the
real entrained pressure and temperature are compared to the entrained pressure and
temperature obtained in a reversible process. Having established a basis for the
definition of such processes, further permutations are left to the reader bearing in
mind that the choice of reference process must be guided by the intended application
of the ejector.
3.2.3 Interpretation of a Thermodynamically Reversible Ejec-
tor Process
Here a thermodynamically reversible ejector process is further elucidated. Impor-
tantly, this illustration does not attempt to explain how a reversible process could
be achieved in a real steady-flow ejector. Rather, it points out how, theoretically, a
reversible process could take place within the control volume of an ejector. For sim-
plicity, let us consider an ejector process with pure fluids at the inlet, both of identical
chemical composition. The a reversible ejector process can now be interpreted as the
sum of two processes. The first process establishes mechanical equilibrium and the
second establishes thermal equilibrium. The process that establishes mechanical equi-
librium at the discharge pressure, PD is in fact a turbine-compressor processs. The
motive and entrained fluids are respectively expanded and compressed through an
adiabatic and isentropic turbine and compressor, Fig. 3-6.
The thermal equilibrating process may be represented by a reversible heat engine that
transfers heat (and entropy) from the the hotter of the fluids M' or E' to the colder
fluid, allowing them to reach an equilibrium temperature TD. Of course, work is
produced by the heat engine to supply the compression process. It is therefore clear,
that in a reversible ejector process, the work available for compression is greater
than that available in a turbine-compressor process. Consequently, the reversible
entrainment ratio will always be greater than or equal to the turbine-compressor
entrainment ratio. In the case where the fluid exiting the turbine and the compressor
in Fig. 3-6 are at the same temperature, no work will be done by the reversible heat
engine. Consequently, in this special case, the reversible entrainment ratio and the
Figure 3-6: Schematic diagram of a thermodynamically reversible ejector for fluids ofidentical chemical composition
turbine-compressor entrainment ratio will be identical. The equilbrium process paths
of Fig. 3-6 are represented in Fig. 3-7.
Until this stage, processes involving only pure and identical inlet fluids have been
considered. However, the same analysis can be generalised to cover processes such as
the compression of a moist air stream using high pressure steam. In such cases, ad-
ditional variables, such as humidity ratio, are required to fully define the equilibrium
state at any point.
3.3 Reversible Entrainment Ratio Calculations for
Different Fluids
At present, the performance of ejectors is typically predicted using graphical meth-
ods or by semi-empirical correlations [28, 29]. These methods allow the calculation of
the entrainment ratio as a function of inlet and outlet fluids' thermodynamic states.
Whilst the reversible entrainment ratio does not provide much information about the
entrainment ratio of real devices, it does provide an upper bound upon the entrain-
M P
E'
D
M' E
Specific Entropy [J/kg K] S
Figure 3-7: Process paths for a thermodynamically reversible ejector
ment ratio achievable. In this section, trends in the reversible entrainment ratio with
inlet fluid conditions and discharge pressure are illustrated for ideal gases (air-air),
steam-steam and steam-moist-carrier-gases ejectors.
3.3.1 Air-air
Derivation of the Reversible Entrainment Ratio for Ideal Gas Ejectors
For two reasons, the ideal gas ejector model is central to the understanding of ejec-
tor operation and performance characterization. First, the ideal gas assumptions
allow analytical expressions to be derived for ejector performance. Secondly, for per-
formance characterization of ejectors, the ideal gas model is appropriate for many
ejector processes, as will be seen in Section 3.3.2.
Here an analytical expression is derived for the reversible entrainment ratio of an
ideal gas process. The specific heat capacities of the gases at constant pressure are
assumed to be equal and constant. The equation of state for an ideal gas is:
Pv = RT (3.9)
where P is the total pressure, v the specific volume, R the ideal gas constant on a
unit mass basis and T the total temperature. Equation (3.2) for ideal gases takes the
following form for a reversible ejector process:
cTm + ER - cpTE = (1 + ER)cpTD (3.10)
Rearranging this equation, the discharge temperature may be written in terms of the
inlet conditions and the reversible entrainment ratio:
ER -TETM (3.11)1 +ER
Rearranging Eq. (3.11) the following dimensionless equations are obtained:
TM 1+±ERTD 1+ER (3.12)TD 1 + ER - ITYTM
Tm- + ERTD = E (3.13)TE 1 +ER
Next, rearranging Eq. (3.7), the Second Law may be written as follows:
SM -r Sgen
ER = S V + mM (3.14)(Srev SE)
Here we seek an expression linking the change in entropy to the state properties T
and P. The following expression describes a differential change in entropy of an ideal
gas:
dT Pds = T- T dP (3.15)
dT dPds = cP R- (3.16)
Now the entrainment ratio may be written as follows:
+gen f M D M
ER = D d D ? (3-17)fE % -T - E --
For constant specific heats Eq. (3.17) results in the following relation describing the
reversible entrainment ratio:
Sgen T Rl- R -"+ cpln(.L) -RnN
ER = mm TD "D (3.18)cyln(g) - Rln(-2)
Dividing the right hand side of Eq. (3.18) by the ideal gas constant, R, renders
all terms dimensionless. Note also the definition, purely for convenience, of the di-
mensionless entropy generation rate.
s gen + -'t1-In (T-M- -][ n( )ER = g ± ( TD - (3.19)
-In -T) -lIn(fa)-1 TE' O
__ genSgen - r*" (3.20)
rnMR
Eq. (3.12), (3.13) and Eq. (3.19) allow the entrainment ratio to be expressed as a
function of the following dimensionless parameters:
ER = fn(TE/TM, PE/PM, PD/PM,'7,gen) (3.21)
The reversible entrainment ratio, RER, can now be defined in terms of the inlet
fluid states and the discharge pressure using Eq. (3.19).
In(TA) - In (P)RER = D1 TD
-7ln(-) - ln( )-1 TE _
3.3.2 Steam-steam Reversible Entrainment Ratios
In this section, the reversible entrainment ratio is calculated for steam-steam ejectors.
As a first step, let us ask how well steam-steam ejectors can be modelled using ideal
4000- 400P= 100bar 'P=0bar ,P =I bar
03500 -
3000-
2500-
Point
2000 '
/ '' ' ,' P=0.Ibar
-/ y1.29
4 5 6 7 8 9 10
Specific Entropy (kJ/kg-K)
Figure 3-8: Lines of constant ratio of specific heats, -y, for steam
gas approximations. We note that steam-steam ejectors are typically supplied with
supersaturated motive steam. This is done to avoid condensation within the ejector
and a subsequent loss of mass via a condensate stream.
Steam behaves like an ideal gas at conditions of high temperature or low pressure.
Supersaturated steam is very well approximated by ideal gas assumptions at pressures
below approximately 10 bar. Above this pressure, ideal gas assumptions are only
valid at higher levels of superheat. In Fig. 3-8, the saturation curve of pure water is
plotted on axes of specific enthalpy versus specific entropy. The ratio of specific heats,
-y, is calculated at a variety of points within the superheated region and plotted to
demonstrate that 7 is approximately constant in the low pressure superheated zone.
The entrained pressure and discharge pressure for steam ejectors are typically
close to or below atmospheric pressure. Therefore, if the motive steam pressure is
below about 10 bar, Eq. (3.12), (3.13) and (3.21) are approximately applicable in
describing the inlet and outlet states of typical steam-steam ejectors.
The exact calculation of RER for steam is relatively straightforward. In each case
the equations solved are Eq. (3.6) & (3.7). (If the state lies under the saturation
curve, T and P would be insufficient to determine h and s. Instead, T and x, the
3000
2 bar2800 -
M ia(ho", sI Dar
2600co
C2400
0
2200 parallelCritical
Point
20004 4.5 5 5.5 6 6.5 7 7.5
Specific Entropy (kJ/kg-K)
Figure 3-9: Specific enthalpy-entropy diagram for pure water
dryness fraction, or P and x would be required.) To graphically explain trends in the
reversible entrainment ratio of a steam-steam ejector, Fig. 3-9 is provided. Figure 3-9
shows the states of the motive, discharge and entrained fluids on a specific enthalpy
versus specific entropy diagram. The state of the discharge fluid is given by the
intersection of the isobar at the discharge pressure and a line joining the motive
and entrained states (again see Sect. 3.2 for derivation). The entrainment ratio is
represented by the length of the line from the motive to the discharge state, |MDI,
divided by the length of the line from the discharge to the entrained inlet state, IDE|,
Eq. (3.23).
RER = MD (3.23)|DE
For a fixed state of the entrained fluid, it may be shown that a motive saturated
steam pressure exists that maximizes RER. In Fig. 3-10, RER is plotted as a func-
tion of the motive steam pressure, choosing inlet fluid dryness fractions of 100% for
illustrative purposes. RER increases logarithmically until a motive steam pressure of
approximately 15 bar is reached. Beyond a motive saturated steam pressure of 15
612
.6 8
00 50 100 150 200
Motive Steam Pressure, Pm [bar]
Figure 3-10: Reversible entrainment ratio versus motive steam pressure for a steam-steam ejector. (PD=15 kPa, PE=10 kPa, XM=1, XE=1)
bar, the rate of increase in RER reduces until an optimal motive steam pressure is
reached. This optimum pressure corresponds to the point at which the tangent to
the saturation curve is parallel to the slope of the isobar at the discharge state (see
Fig. 3-9). This conclusion is reached by considering the point at which the distance
|MDI divided by IDEl is maximised subject to point D lying at the intersection of the
isobar at the discharge pressure and a line joining states M and E. The optimal pres-
sure of the saturated motive fluid therefore depends upon the discharge pressure. As
the discharge pressure changes, the slope of the isobar changes and thus the point at
which the tangent to the saturation curve is tangential to the isobar at the discharge
pressure also changes.
In Fig. 3-11, the optimal motive steam pressure is plotted for a fixed entrained
fluid state and discharge pressure. The results are primarily of academic interest
given that the optimal motive steam pressures are well beyond the typical pressures
used in practical applications.
In Fig. 3-12 RER is plotted as a function of the discharge steam pressure. It is
seen that RER drops drastically with an increasing compression ratio. Often, the
130 r
125
2 120
0 115
2 110
o 105
1000 20 40 60 80 100 120
Discharged Pressure, PD [kPa]
Figure 3-11: Optimal motive steam pressure of a steam-steam ejector
motive steam may be superheated and therefore it is interesting to consider the effect
of the degree of superheating on RER. Figure 3-13 highlights the improvement in
RER due to superheating at constant pressure of the motive steam.
In contrast to Fig. 3-10 and Fig. 3-12, the rate of change in RER is small with
changes in the inlet motive steam temperature. Again, this trend may be visualised
making use of Fig. 3-9. As the motive steam is superheated following an isobaric pro-
cess, both its specific enthalpy and its specific entropy increase, although its specific
enthalpy increases at a greater rate. This results in an increasing reversible entrain-
ment ratio. The apparent linearity of Fig. 3-13 may be explained by the fact that the
isobars in the superheated region of Fig. 3-9 are each approximately linear. Thus, the
increase in reversible entrainment ratio, according to Eq. (3.23), will be linear.
Finally, we consider how RER can be maximized when there is an upper limit on
the motive steam temperature. To do this we may plot RER at constant temperature
against motive steam specific entropy, as is done in Fig. 3-14.
The line plotted in Fig. 3-14 is a line of constant temperature. The reversible
entrainment ratio reaches a maximum when x=1, i.e. when the steam is in a saturated
condition. Consider the entrainment of saturated steam at state E by saturated
Compression Ratio, 140 2 4 6 8
50
,40
030
e20
1-4 100
0' ' ' '
10 12
0 20 40 60 80 100
Discharged Pressure, PD [kPa]
Figure 3-12: Reversible entrainment ratio versus discharge pressure for a steam-steamejector. (PM=10 bar, PE=10 kPa, xM=l, ZE=1)
16
14
12
10
8
6
'$4
e2
00 25 50 75 100 12
Motive Steam Superheat, Tm - Tw [*C]
Figure 3-13: Reversible entrainment ratio versus motive steam superheat for a steam-
steam ejector. (PD=15 kPa, PM= 10 bar, PE=10 kPa, XE=l, TM,sat = 180'C
10
8-
6
4
Saturation Superheat2 - XM<1 1XM
02 3 4 5 6 7 8
Motive Steam Specific Entropy, sm [kJ/kg-K]
Figure 3-14: Reversible entrainment ratio versus motive steam specific entropy fora steam-steam ejector, with motive steam at constant temperature. (TM=179.9*C,Psat,M=10 bar, PD=15 kPa, PE= 10 kPa, XE=1)
motive steam at state M (see Fig. 3-15). If the condition of the motive steam is
brought below the saturation curve, the ratio of the line IMDI to |DE decreases.
Thus, because the isobars on a h - s diagram of steam converge as the state moves
from that of saturated steam to that of a wet mixture, RER using saturated motive
steam will always be superior to that using a wet mixture at the same temperature. In
the superheated region of the h - s diagram, isobars and isotherms do not coincide.
If state M moves along an isotherm into the superheated region, RER is severely
compromised, as the isotherm corresponding to state M converges rapidly with and
meets the isobar corresponding to the discharge state D. In real steam-steam ejectors,
slightly superheated motive steam may preferentially be used to saturated steam in
order to discourage the formation of droplets in the discharge, which may need to be
removed using a mist eliminator.
As a final caveat, maximizing RER does not in any way maximize the exergetic
efficiency of a real device. The RER represents a perfect and reversible process that
is 100% efficient from a Second Law perspective. Rather, the purpose of this section
3000
2800
T=453 K,Pi 0 bar D
S260 ET=393.4 K,Pi=2 bar
2400 -
T=372.8 K,Pi=1 bar
2200CriticalPoint
2000-4 5 6 7 8
Specific Entropy (kJ/kg-K)
Figure 3-15: Isotherms on a specific enthalpy-entropy diagram for pure water
was to comprehend the upper thermodynamic limit on the entrainment ratio.
3.3.3 Steam-Moist-Gas Reversible Entrainment Ratios
Multipressure humidification-dehumidification (HDH) desalination cycles, which pre-
viously have been proposed by the current authors, make use of ejectors to pressurize
a moist carrier gas [4, 5, 30]. It is therefore of interest to consider the effect upon
the reversible entrainment ratio, RER, of using carrier gases of different types. In the
case of a steam-moist-carrier-gas ejector, the entrainment ratio is defined as follows:
E _(1 + WE)rhcER =1 E (3.24)M
where WE is the humidity ratio of the entrained fluid. In the case of humidification-
dehumidification desalination, the quantity of water vapor entrained is of importance,
rather than the total amount of fluid entrained. For this reason, it is useful to define
the vapor entrainment ratio. Of course, the vapor entrainment ratio is equivalent on
a mass and molar basis as long as the motive fluid is steam.
ERa WEmc WEmyap 1+RWE (3.25)
In this section the entrained and discharge fluids are treated as ideal mixtures
of ideal gases, within which the enthalpy and entropy of the carrier gas and steam
are calculated at their partial pressures. The ratio of the number of moles of carrier
gas and water vapor present in a moist carrier gas mixture along with the relative
humidity are given as follows:
Pvap Nst _ Psat
P -Pvap Ncg P - 4Psat(3.26)
When comparing the reversible entrainment ratio of different moist carrier gases, we
do so at equal entrained fluid temperature, pressure and relative humidity. Therefore,
it is sensible to consider the governing equations on a molar basis, since ( is constant
for all carrier gases under these conditions.
First Law of Thermodynamics (reversible process):
NSt,Mhst,M + Ncghcg,E + Nst,Ehst,E = Nst,DhrtD + NcghcgD (3.27)
where N is a molar flow rate and h is a molar enthalpy. According to a mole balance
on steam:
NSt,DRERva
ht,m + vah Icg,E + RERvaphst,E(E
= Nt,M + st,E (328)
= (1 + RERrap)hDev grev (.9=D ±E cgD
Second law of Thermodynamics (reversible process):
Nst,'MgstM + Ncgscg,E + Nst,Esst,E = Nst,D stD NcgcD (3-30)
Sst,M + RE gcg,E + RERvapSst,E -(331)
(1 + RERvap)gev + RERvap gev 3.32)stD + E cg,D
Equations of State (reversible process):
S= fn(prev rev (3.33)
*v = f(prev rev (334)
§ is the molar enthalpy. The partial pressures of the steam, Pst,D, and carrier gas,
Pcg,D, both sum to the total discharge pressure, PD. RER and RERVap are now plotted
in Fig. 3-16 and Fig. 3-17 as a function of motive steam pressure, setting zM and #Eto unity for illustrative purposes. 2
The results of Fig. 3-16 collapse onto a single curve in Fig. 3-17. Surprisingly, RERyap
appears to be almost independent of carrier gas. Finally, in Fig. 3-18 the reversible
vapor entrainment ratio is plotted versus the entrained fluid inlet temperature, main-
taining a relative humidity of unity.
The increase in RERvap with temperature of the entrained moist carrier gas is
explained by the decrease in value of the specific heat of the entrained fluid per mole
of vapor.
3.4 Development of Ejector Performance Metrics
When one seeks to evaluate the thermodynamic ideality of a process, performance
metrics are often chosen that compare useful work done in a real process to that in a
defined reference process. However, the identification of work done within an ejector
2Note: Although not graphed here, the variation of reversible entrainment ratio with compressionratio for helium, air and carbon dioxide carrier gases is qualitatively similar to that obtained forsteam-steam ejectors.
30 -
25 -
20 -
15 -
10
00 50 100 150 200
Motive Steam Pressure, Pm [bar]
Figure 3-16: Reversible entrainment ratio versus motive steam pressure for steam-moist-carrier-gas ejectors. (PD=75 kPa, PE= 5 0 kPa, TE=50 *C, xm=l, qt =1)
4
3.5
3
2.50.
2
1.5
1
0.5
0
-Carbon Dioxide
- Air
- -Helium
0 50 100 150
Motive Steam Pressure, PM [bar]
Figure 3-17: Reversible vapor entrainment ratio versus motive steam pressure forsteam-moist-carrier-gas ejectors. (Ilc=1.5, PE= 50 kPa, TE=50 *C, xM=l, qE=1)
5
4
3
2--Carbon Dioxide
--Air
-- Helium
010 20 30 40 50 60 70 80
Entrained Carrier Gas Temperature, TE [*C]
Figure 3-18: Reversible vapor entrainment ratio versus entrained saturated carrier gastemperature for a steam-moist-carrier-gas ejector. (PD=75 kPa, PE= 50 kPa, IIi=20,XMz=1, =1)
is not a trivial matter. In this section, these dilemmas are discussed and suggestions
are made as to how performance metrics can be defined.
The crux of difficulties in evaluating work done within an ejector is that one
cannot easily distinguish between useful flow work, heat transfer and dissipation. A
similar difficulty is very clearly explained by Casey and Fesich [31] in relation to
diabatic turbochargers, where heat transfer between the compressor and the turbine
is significant. Maintaining the nomenclature of Casey and Fesich, the Gibbs equation
may be integrated along the process path of a single fluid stream:
h2- hi= jv dp + j(T ds)q + j(T ds) irrev = Y12 + q1 2 + 12 (3.35)
where Y12 is the useful flow work done in the process, q12 is the reversible heat trans-
fer and i12 is the dissipation, or in other words, the lost work. Whilst the change
in enthalpy of the fluid is fixed by the end states, the useful flow work, the heat
transferred and the dissipation are all path dependent. Only in the special case of an
adiabatic and reversible process can the useful flow work be equated with the change
in enthalpy of the fluid:
h2 - hi = v dp (3.36)1
In reality, the motive and entrained streams in an ejector remain neither adiabatic
nor distinct throughout the process. In real ejector processes, both heat transfer and
dissipation are experienced by the motive and entrained streams during mixing. Con-
sequently, the useful flow work can only be evaluated if the process path is known.
Unfortunately, ejector mixing processes are characterized by a high degree of macro-
scopic disorder, involving heat transfer with finite temperature differences, forces that
are not fully restrained and irreversible mixing. Intermediate states cannot accurately
be represented on a process path of one intensive quantity against another. Conse-
quently, there is no process path or equation in the form of Eq. (3.35) that can be
applied to actual ejector processes.
Despite the difficulty of identifying useful work done in an ejector, there remain
two broad approaches to characterising performance:
1. Comparisons may be made between terms representing useful work done in
real and reference processes, provided that a clear rationale is given as to why
such a metric compares different systems at different operating points. (Some
performance metrics may only evaluate the performance of a part of the entire
ejector process, as shall be seen with the turbine-compressor or compression
efficiency.)
2. An exergy based approach may be taken. A performance metric may be defined
once the useful exergetic output, the exergetic input, and a reference state for
the process have been decided upon.
Next, four efficiencies based on the comparison of real and reversible processes are
presented, followed by an exergetic efficiency.
3.4.1 Reversible Entrainment Ratio Efficiency
The definition of reversible entrainment ratio efficiency was provided in work by
Elrod [19]. Al-Khalidy et al. [27] referenced the work of Elrod and also employed
this efficiency. The reversible entrainment ratio efficiency r/RER quantifies how close
the actual entrainment ratio of an ejector is to the theoretical maximum, for fixed
conditions of operation. Its value is therefore always between zero and unity. RER is
calculated based upon the reversible process described in Section 3.2.2.
ERr/RER = RER (3.37)
For a process involving fluids with identical ratios of specific heats and behaving
as ideal gases, the RER may be derived based upon the ideal gas analysis conducted
in Section 3.3.1. The ratio of the real to reversible entrainment ratio may be written
as follows:
S(en + In( ) -n( ) ln( ) - In(PM)ER -y - 1 TD PD 7y - 1 D 3D
RER 7 ln(-) - I( )M ln( I - MIn(- - 1 TD PD7 -y TD D
One interesting case of ejector operation to consider is where the inlet temper-
atures of both ideal gases are equal. According to the First Law, the discharge
temperature must equal the inlet temperatures in both real and reversible cases.
Equation (3.22) then simplifies to a ratio of the natural logarithm of the expansion
ratio PM/PD to the ratio of the natural logarithm of the compression ratio PD/PE.
PM_ln(-)
RER - PD (3.39)ln(PD)
PE
and the ratio of ER to RER simplifies to:
PMER ln(-) - 9gen
=PD (3.40)RER PM
ln(-)PD
Finally, as a general comment, the principles of operation of steady flow ejectors
impose further limits upon ejector performance that are not captured by a control
volume analysis whereby the control volume surrounds the ejector. These internal
limits depend both upon ejector geometry and operational conditions. Some of the
most enlightening work in this regard has been conducted by Chou et al. [21] and
Arbel et al. [22]. The same is true for the reversible discharge pressure efficiency, to
be presented in the following section.
3.4.2 Reversible Discharge Pressure Efficiency
The reversible discharge pressure efficiency quantifies how close the actual rise in
pressure of the entrained fluid is to the theoretical maximum pressure rise. As with
the reversible entrainment ratio efficiency, it is bounded between zero and unity. The
theoretical rise in pressure is calculated based upon Section 3.2.2. This efficiency is
introduced and discussed in detail by Arbel et al [22].
PD -PEr = PD(3.41)
P~v PE
3.4.3 Turbine-Compressor Efficiency
The turbine-compressor efficiency [1, 32-34] (also known as the turbomachinery ana-
logue efficiency), allows the comparison of an ejector with a mechanically coupled
turbine and compressor. The efficiency is defined as the ratio of the entrainment
ratio in a real ejector to the entrainment ratio achieved in a turbo-charger exhausting
to the same discharge pressure, PD, as the ejector, and within which the turbine and
compressor are adiabatic and isentropic (Fig. 3-19).
h
M>1 E'
4) ECI'I
Specific Entropy [J/kg-K] S
Figure 3-19: Isentropic expansion and compression processes
ERT7ER =~TER (3.42)
Here, TER stands for Turbo Entrainment Ratio or Turbine-Compressor Entrainment
Ratio. There is an important difference between the reversible entrainment ratio
RER and the turbo entrainment ratio TER. The RER is calculated based on a pro-
cess whereby the discharge is a single stream homogeneous in total temperature, total
pressure and chemical composition. The TER is based upon a process whereby the
inlet streams remain distinct at the discharge, i.e. the streams are not required to
be in thermal or chemical equilibrium. Consequently, the value of TER will always
be less than or equal to RER. A corollary of this is that YTER is not bounded by an
upper limit of 100%, as a real ejector could be imagined that achieves an entrainment
ratio greater than that of a turbo, through the exploitation of the thermal or chemical
disequilibrium of the fluids. Of course, a real steady-flow ejector could not in practice
achieve a turbine-compressor entrainment efficiency beyond 100%. There is no obvi-
ous means by which a disequilibrium in chemical potential between the streams may
be exploited. The question as to how thermal disequilibria may be exploited is more
complex. As shall be seen in Sect. 3.5, for ejectors designed to date, the difference in
the value of 7 7TER and 77RER is small. Finally, 7TER may also be computed for a system
comprising an independent turbine and compressor with the role of expanding one
fluid and compressing a second, if the isentropic (also known as adiabatic) efficiencies
of the turbine and compressor are known. Bulusu et al. [1] demonstrates the following
equality:
ER7TER = TER = "'turb,'comps (3.43)
where 7iturb,s and 7eomp,, are the isentropic efficiencies of the turbine and compressor.
Consequently, it may be said that an ejector with a turbine-compressor efficiency of
35% will provide the same performance as a turbine and compressor with isentropic
efficiencies of 50% and 70% respectively, for example.
3.4.4 Compression Efficiency
Al-Najem et al. [23] presented the following measure of ejector performance, which
here is denoted the compression efficiency:
= (lhM + rhE) (hD - hE)rnM(hM - hM,NE)
where hM,NE represents the static enthalpy at the exit of the motive fluid nozzle.
Al-Najem et al. mention that this efficiency compares the compression work done in
the ejector to the kinetic energy available from the motive stream. In a steady flow
ejector, the nozzle through which the motive fluid is accelerated is adiabatic. The
quantity hM - hM,NE represents the kinetic energy gained by the motive flow from the
nozzle inlet to the exit. In an adiabatic mechanical compressor, the compression work
done is given by AhE(hD - hE). Based upon this assertion, it seems likely that the
quantity hD - hE was selected to represent the compression work done in an ejector.
However, in an ejector, heat transfer occurs between the motive and entrained streams
during mixing. Therefore, the change in enthalpy hD - hE is due to a combination
of heating/cooling and work. Indeed, the useful compression work done in an ejector
requires the thermodynamic states describing the process path of the entrained fluid
to be known. In conclusion, based upon the review of literature conducted, the
rationale behind and the utility of the compression efficiency is unclear.
3.4.5 Exergetic Efficiency
Exergetic analysis is a means of evaluating ejector performance from a Second Law
point of view. The premise of exergetic efficiency is to compare the useful exergetic
output of a system or component to the exergetic input. First, for an exergetic
efficiency to be meaningful, an environmental dead state must be defined. Let us
consider an ejector operating within an environment at To and PO, as in Fig.3-2. QO
represents the rate of heat input from the surroundings. For a steady-state process,
the First and Second Laws take the following form:
(hM - hD,M) ± Q0 = ER (hD,E - hE) (3.45)
(SM - SD,M) ± ± Sgen = ER (SD,E - SE) (3.46)
where states D,M and D,E represent the state of the motive and entrained fluids in the
discharge stream, calculated at their respective partial pressures. If the motive and
entrained fluids do not differ in chemical composition, these states will be identical.
Multiplying the Second Law by To and subtracting the resulting equation from the
First Law, the ejector process may be described in terms of flow exergy, x, a state
variable.
(hM - hD) - TO (SM - SD) = (hD - hE) - T0 (SD - SE) - ToSgen (3.47)
-Aim = ERAiE ± TSgen (3.48)
AiE = (hD - hE) - T0 (SD - SE) (3.49)
AiM = (hD - hM) - T0 (SD - SM) (3.50)
The next step is to identify the exergetic output from and exergetic input to
the system. This choice is somewhat subjective and depends on how the purpose
of the ejector is defined. It may be argued that the role of an ejector is to transfer
exergy from the motive fluid to the entrained fluid. This is somewhat analagous to a
compressor or an extraction fan, where the device's role is to transfer exergy to a fluid.
The value (or output) of the ejector process, in exergetic terms, may be measured by
the change in exergy of the entrained fluid. The price paid to implement the process,
again in exergetic terms, is the change in exergy of the motive fluid. Hence, the
following formulation for exergetic efficiency, identical to that employed by Al-Najem
et al. [23], is obtained:
ERx = (3.51)
Considering Eq. (3.47) it is clear that ?1x also quantifies the exergetic losses per unit
of exergetic input to the ejector process, Eq. (3.52).
- - ~ ToSge1- qx = To "e (3.52)
rhM [(hD - hM) - T0 (SD - SM)]
Comparison of the Exergetic and Reversible Entrainment Ejector Efficien-
cies
An analytical expression for the exergetic efficiency may be obtained when the inlet
fluids are ideal gases of identical and constant specific heats. If the further restriction
of having inlet fluids at the same temperature is imposed, the discharge enthalpy
becomes independent of the entrainment ratio, Eq. (3.11). The discharge temperature
equals the inlet temperature, and thus outlet and inlet specific enthalpies are also
equal, Eq. (3.10). Consequently, the exergetic ejector efficiency takes the following
form.
ER(hD - TOSD) ~ (hE -TOSE)(hM - TosM) - (hD - TOSD)
= ER TOSD+TOSE (3.54)-TSM + TSD
= ER SD - SE (3.55)SM - SD
PDin-= ER E (3.56)
PnMPD
From Eq. (3.18), we an obtain the following expression for ER:
PM genln(T-) - S"ER= PD rhM (3.57)
ln(P)
Upon substituting Eq. (3.57) into Eq. (3.56), the exergetic ejector efficiency is found
to be equal to the reversible entrainment ratio for an ideal gas ejector with identical
specific heats and identical inlet temperatures:
In( PM) Sen
x= n P -M RrhM = ?7 RER (3.58)
ln(-)InED
3.5 Comparison of Ejector Performance Metrics
using Experimental Data
Now, having rigourously defined measures of ejector performance we may put these
measures of efficiency to use in order to answer the following questions:
1. How do the numerical values of these efficiencies compare to one another?
2. What are the values of efficiency achieved by devices documented in literature?
40
-RER10 - RDP
01 1.1 1.2 1.3 1.4 1.5 1.6
Compression Ratio, 1, [-]
Figure 3-20: Comparison of efficiency definitions for a steam-steam ejector over arange of compression ratios
To answer the first question, we can simply compare a sample of fictitious ejector per-
formance data. For example, we can compare the disparity in efficiency definitions
over a range of compression ratios, PD/PE. Let us consider a single case involving
the compression of saturated water vapor using high pressure saturated steam. Let
us choose PM= 10 bar and PE= 5 0 kPa. Let us fix the performance of the real ejec-
tor by fixing ER (Note: we could also theoretically fix the real outlet state or the
dimensionless entropy generation sgeen). Now, in Fig. 3-20 the range of efficiencies are
plotted versus the compression ratio, PD/PE- The compression efficiency is omitted
due to its lack of rigour and its requirement for knowledge of a thermodynamic state
internal to the ejector control volume.
To answer our second question, we must extract data from literature. One draw-
back of using such data is that the performance of ejectors analysed in literature is
not necessarily representative of the state of the art ejectors employed in industry.
Furthermore laboratory tests may not have been performed at the ejectors' design
points of operation. With this caveat in mind, the efficiencies of an air-air and a
steam-steam ejector are analysed, acknowledging that the conclusions drawn may
16
14 -
12 -
#10 - 'RDP
8 *TER
oAX6 - RER
+ A A4~ A A
2A
00 1 2 3 4 5
Compression Ratio, 1, [-l
Figure 3-21: Comparison of efficiency definitions using data from an air-air ejector,[1, 2]. Chamber throat diameter, 0.88 inches; Diffuser exit diameter, 3.02 inches;Nozzle throat diameter, 0.5625 inches; Nozzle exit area 0.735 inches; Plenum diameterfor secondary flow, 3.25 inches.
not be generalised to cover all ejectors. Amenable data for air-moist-carrier-gas ejec-
tors was not available. Data for the air-air ejector was extracted from work by Bulusu
et al. [1], Fig. 3-21. Data for the steam-steam ejector is from work by Eames et al.
[3], Fig. 3-22. One should be aware that the degree of superheat of the motive steam
is not obvious from the reported work and motive steam has been assumed to be
saturated at the entrance to the ejector.
Note how, according to Eq. (3.58), the exergetic and reversible entrainment ratios
are identical in value. The discrepancy in the turbine-compressor efficiency compared
to the former two definitions is due to the fact that entropy generated in reaching
thermal equilibrium of the fluids is unaccounted for.
Note that critical conditions of operation typically refer to the mode of ejector
operation where the discharge pressure is such that choking of the entrained fluid
occurs within the ejector. This choking is due to the formation of a hypothetical
throat for the entrained fluid by the motive fluid expanding from the nozzle. For
given inlet fluid conditions, the critical point of operation corresponds to the point
20 -
*TER-15 - *
A U 0ORER
.RDP e10 -
*e5 so *
00 0.1 0.2 0.3 0.4 0.5 0.6 0.7
Entrainment Ratio, ER [-]
Figure 3-22: Comparison of efficiency definitions using data from a steam-steamejector operating at conditions of critical back pressure [3]. Chamber throat diameter,18 mm; Diffuser exit diameter, 40 mm; Nozzle throat diameter, 2 mm; Nozzle exitarea 8 mm; Plenum diameter for secondary flow, 24 mm.
where the entrained mass flow rate is maximum. The occurance of choking within
steady flow ejectors is dealt with in great detail by Chou et al. [21]. In Fig. 3-22
as in Fig. 3-21, the exergetic and reversible entrainment ratio efficiencies are similar
in value although both demonstrate an increasing trend. The turbine-compressor
efficiency is also close, and as expected, is greatest in value. Meanwhile the reversible
discharge pressure efficiency differs significantly in value. This may be explained
in two ways. Firstly, the reversible processes employed for the reversible discharge
pressure and the reversible entrainment ratio are different. Secondly, the ratio of mass
flows, employed in the reversible entrainment ratio, and the ratio of pressure rises of
the entrained fluid, employed in the reversible discharge pressure clearly do not scale
in the same manner.
3.6 Conclusions
In this chapter, various definitions of ejector efficiency have been evaluated. When
defining or employing an ejector efficiency, it is crucial to describe the physical sig-
nificance of the particular efficiency along with its utility. The following conclusions
have been reached:
1. A reversible entrainment ratio may be calculated for any steady-flow ejector
process, by employing the the First and Second Laws of thermodynamics, pro-
vided the thermodynamic states of the inlet fluids and the total pressure of
the discharge fluid, all at equilibrium conditions, are known for a real ejector
process.
2. The reversible entrainment ratio of an ideal gas ejector may be formulated
analytically as a function of the inlet (motive-to-entrained) temperature and
pressure ratios, the compression (discharge-to-entrained) pressure ratio and the
ratio of specific heats. Calculation of the real entrainment ratio also requires
knowledge of a further state variable describing the discharge (or alternatively
the dimensionless entropy generation).
3. For steam-steam ejector applications, the inlet and outlet ejector states are
shown to be accurately represented by ideal gas equations for pressures below
10 bar. Higher pressure motive steam may also be modelled as an ideal gas,
provided the level of superheat is sufficient.
4. Where the motive fluid employed in a ejector is saturated steam, there exists
an optimal pressure that maximizes the reversible entrainment ratio.
5. Where the temperature of motive steam is limited, the reversible entrainment
ratio is maximized by employing saturated motive steam at a pressure equal to
the saturation pressure at the maximum allowable temperature.
6. The reversible vapor entrainment ratio (ratio of vapor entrained to motive va-
por) for steam-moist-carrier gas ejectors is insensitive to the choice of carrier
gas, if the inlet relative humidity is maintained constant.
7. The turbine-compressor entrainment ratio differs from the reversible entrain-
ment ratio only in that the discharge fluids are unmixed and not brought into
thermal and chemical equilibrium.
8. The reversible entrainment ratio RER is always greater than or equal to the
turbine-compressor entrainment ratio TER.
9. The reversible entrainment ratio efficiency iRR is always lower than the turbine-
compressor efficiency 7TER-
10. For general air-air and steam-steam ejectors, the exergetic efficiency qx is very
close in numerical value to the reversible entrainment ratio efficiency nRER-
11. The exergetic efficiency qx for an ideal gas ejector with inlet fluids at the same
temperature is identical to the reversible entrainment ratio efficiency /RER.
12. The reversible entrainment ratio efficiency RER is advantageous compared to
the exergetic efficiency qx as its value does not depend upon the choice of a
reference temperature.
Chapter 4
Recommendations
The research outlined in Chapter 2 leaves us with a clear idea of the limitations upon
the performance of HDH systems, particularly where components of large size (small
pinch point temperature and concentration differences) are concerned. However, a
methodology for optimizing systems of fixed size is still required. The ejector research
in Chapter 3 leaves us with a clear understanding of how the performance of different
ejectors can be compared. In addition, we have a clear idea of how reversible ejectors
would behave and can therefore quantify the limits upon real ejector performance.
The next research step would be to investigate exactly why and when real ejector
performance approaches reversibility and why and when this is not possible.
As is often the case, when we grow in understanding, opportunities to apply this
understanding in the form of a new or improved technology arise. Here, two promising
research ideas discovered during the course of my research, are outlined.
4.1 Reduced Temperature Range HDH Cycles
One of the key insights into the performance of HDH systems in Chapter 2 relates to
our new understanding of the effect of temperature range upon performance. It was
established that by reducing the temperature range of an HDH cycle, the GOR greatly
improves whilst the recovery ratio decreases. One means of improving GOR would be
to reduce the top temperature of the HDH cycle. Unfortunately, a lower temperature
, |HXTr 'r
Figure 4-1: Low temperature range HDH configuration
HDH system operating at atmospheric pressure would also suffer from low fractions
of vapor throughout, and conversely high carrier gas fractions. This higher fraction
of carrier gases would lead to increased heat and mass transfer resistances, thus
increasing the size and cost of the system.
Another, less obvious, solution is to increase the bottom temperature of the cycle.
This may be done by employing a closed air cycle with a regenerating heat exchanger
for the feed and reject streams, as in Fig. 4-1. Not only would the reduced temperature
range improve the GOR, but the high fraction of vapor at higher temperatures would
facilitate improved heat and mass transfer coefficients, leading to a smaller system
size.
4.2 Future Research Directions
For heat and mass exchangers of finite size, the question remains as to what the
optimal top and bottom temperatures would be. Furthermore, if the carrier gas is to
operate in a closed loop, the choice of a carrier gas other than air may be advantageous,
100
particularly for reasons of improved thermal conductivity. Such questions would build
upon the work of Narayan et al. [35].
4.3 Ejector-Expander Power Generation System
We now return to Fig. 1-1, where the combination of an ejector with an HDH system
was presented in the introduction. This multi-pressure system was investigated in
detail by Narayan et al. [5]. Due to the low amount of water produced in the dehu-
midifier of such a system relative to the power output of the expander, the system was
found to be a power production cycle. Following this train of thought, and removing
the humidifier from the system, the power production system of Fig. 4-2 is obtained.
Pure water flows in the upper loop and is pumped to high pressure, converted
from the liquid to vapor phase in the heater, expanded within the ejector and then
condensed in the dehumidifier. A carrier gas would flow in the lower loop, being
compressed in the ejector, dehumidified and then expanded within the expander. The
use of a bubbling column dehumidifier as investigated by Narayan et al. [36] may be
ideal for such an application. Questions remain as to what the ideal pressures would
be to achieve an optimal blend of efficiency and low capital cost. Furthermore, since
the cycle operates with two closed loops, there is significant flexibility in the choice
of fluids for both the upper and lower loops. The development of higher efficiency
ejectors would also be a crucial step.
101
Ts| Ejector
lhda III4 ma ,B
Pahigh PUMP
APDH
PH PDH
Power OutExpander
Figure 4-2: Ejector-expander shaft-power generation system
102
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